Seismic Vulnerability Assessment and Retrofitting Strategies for Masonry Infilled Frame Building (Springer Theses) 9783031203718, 3031203712

Table of contents :
Supervisor’s Foreword
Parts of this thesis have been published in the following documents
Acknowledgements
Contents
1 Introduction
1.1 Framework and Motivation
1.2 Objectives and Tasks
1.3 Thesis Organization
References
2 Out−of−Plane Behaviour of Masonry Infill Walls Experimental Tests: A Systematic Review
2.1 Introduction
2.2 Methodology
2.3 Selection Criteria
2.4 Study Selection
2.5 Data Extraction, Quality Assessment and Data Synthesis
2.6 Final Database
2.7 Global Overview of OOP Tests on Infill Masonry Walls
2.8 OOP Load Application Strategies and Protocol
2.9 OOP Behaviour of As-Built Infill Masonry Walls
2.9.1 Effect of RC Frame Type and Detailing
2.9.2 Effect of Geometric Dimensions and Mechanical Properties
2.9.3 Effect of Gravity Load
2.9.4 Effect of Openings and Panel Support Conditions
2.10 Assessment of the Previous IP Damage Effect on the Infill Walls OOP Capacity
2.10.1 Background
2.10.2 Cracking Strength and Secant Cracking Stiffness
2.10.3 Maximum Strength and Secant Stiffness
2.10.4 IP−OOP Ultimate Displacement
2.11 Final Considerations
References
3 Strengthening Strategies to Improve the Seismic Behaviour of Infill Masonry Walls
3.1 Introduction
3.2 Literature Review of Retrofit and Strengthening Techniques
3.2.1 Disconnection of the Infills from the Structural System
3.2.2 Infill Masonry Walls Retrofitting and Strengthening Techniques
3.2.3 Reinforced Plaster
3.2.4 Bed Joints Reinforcement
3.3 Final Considerations
References
4 Experimental Characterization of the As-Built Masonry Infill Components’ Properties
4.1 Framework
4.2 Stage 1—Material Characterization Tests of Masonry Units and Mortar
4.2.1 Introduction
4.2.2 Compression Strength Tests of Masonry Units
4.2.3 Compression and Flexural Strength Tests of the Mortar
4.3 Stage 2—Mechanical Characterization of the Infill Masonry Wallets
4.3.1 Testing Campaign Overview
4.3.2 Mechanical Properties of Mortar
4.3.3 Compressive Strength Tests
4.3.4 Diagonal Tensile Strength Tests
4.3.5 Flexural Strength Tests
4.3.6 Final Remarks on Stage 2 Tests
4.4 Stage 3—Dynamic Characterization of Infill Panels Through Ambient Vibration Tests
4.4.1 Introduction
4.4.2 Methodology
4.4.3 Laboratory Tests
4.4.4 In-Situ Tests
4.4.5 Final Remarks on Stage 3 Tests
4.5 Final Considerations
References
5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built Masonry Infill Walls
5.1 Introduction
5.2 Specimens Details
5.3 Material Properties and Specimen Construction
5.3.1 Concrete and Reinforcement Steel Bars
5.3.2 Mortar and Masonry Units
5.4 Stage 1—Out-of-Plane Tests Using Airbags
5.4.1 Test Setup
5.4.2 Instrumentation and Loading Protocol
5.4.3 Stage 1: Individual Test Results
5.4.4 Stage 1: Global Results’ Comparison
5.5 Stage 2—Out-of-Plane Tests Using Pneumatic Actuators
5.5.1 Test Setup
5.5.2 Instrumentation and Loading Protocol
5.5.3 Stage 2: Individual Test Results
5.5.4 Stage 2: Global Results Comparison
5.6 Overview of the Summary of Global Results: Critical Analysis
5.6.1 Effect of Columns Axial Load
5.6.2 Effect of Reduction of the Panel Support Width
5.6.3 Effect of Previous Damage
5.6.4 Effect of Plaster
5.6.5 Effect of Workmanship
5.6.6 Effect of Test Setup
5.7 Final Remarks and Major Observations from the Experimental Campaign
References
6 Experimental Assessment of Strengthening Solutions to Prevent the OOP Collapse of Infill Masonry Walls Through Textile Reinforced Mortars
6.1 Introduction
6.2 Flexural Strength Tests in Strengthened Infill Masonry Wallets
6.2.1 Objectives and Specimens’ Description
6.2.2 Specimens Description and Strengthening Process
6.2.3 Test Setup Description and Instrumentation
6.2.4 Experimental Results
6.3 Experimental Assessment of TRM Solutions to Improve the Out-of-Plane Capacity of Full-Scale Infill Masonry Walls
6.3.1 Objectives and Specimens’ Description
6.3.2 Description of the Strengthening Strategies
6.3.3 Material Properties
6.3.4 Test Setup, Instrumentation and Loading Protocol
6.3.5 Experimental Results
6.3.6 Comparison of the Results and Discussion
6.4 Final Considerations
References
7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour
7.1 Introduction
7.2 Brief Overview on Numerical Modelling Approaches: From Simplified Macro-models to Detailed Modelling
7.2.1 Simplified Macro-models
7.2.2 Detailed Micro-modelling Approach
7.3 Development of a Simplified Modelling Approach to Simulate Infill Walls in OpenSees
7.3.1 General Considerations
7.3.2 Modelling of the IP Behaviour
7.3.3 Modelling of the OOP Behaviour
7.4 Numerical Modelling Calibration
7.4.1 In-Plane Behaviour—Panels Without Openings
7.4.2 In-Plane Behaviour—Panels with Openings
7.4.3 OOP Behaviour
7.5 Case Study
7.5.1 Description of the Building
7.5.2 Numerical Modelling Strategy
7.5.3 Material Modelling Properties
7.5.4 Preliminary Results, Natural Frequencies and Vibration Modes
7.5.5 Effect of the Infill Walls in the Seismic Vulnerability Assessment of Rc Buildings Parametric Study 1
7.5.6 Effect of the Infill Panels in the Floor Response Spectra Parametric Study 2
7.5.7 Seismic Vulnerability Assessment of Infilled RC Structures Subjected to Mainshock-Aftershock Sequence Parametric Study 3
7.6 Final Considerations
References
8 Final Conclusions and Future Works
8.1 Conclusions and Major Contributions
8.2 Future Works
Author Biography
Bibliography

Citation preview

Springer Theses Recognizing Outstanding Ph.D. Research

Andre Filipe Furtado

Seismic Vulnerability Assessment and Retrofitting Strategies for Masonry Infilled Frame Building

Springer Theses Recognizing Outstanding Ph.D. Research

Aims and Scope The series “Springer Theses” brings together a selection of the very best Ph.D. theses from around the world and across the physical sciences. Nominated and endorsed by two recognized specialists, each published volume has been selected for its scientific excellence and the high impact of its contents for the pertinent field of research. For greater accessibility to non-specialists, the published versions include an extended introduction, as well as a foreword by the student’s supervisor explaining the special relevance of the work for the field. As a whole, the series will provide a valuable resource both for newcomers to the research fields described, and for other scientists seeking detailed background information on special questions. Finally, it provides an accredited documentation of the valuable contributions made by today’s younger generation of scientists.

Theses may be nominated for publication in this series by heads of department at internationally leading universities or institutes and should fulfill all of the following criteria . They must be written in good English. . The topic should fall within the confines of Chemistry, Physics, Earth Sciences, Engineering and related interdisciplinary fields such as Materials, Nanoscience, Chemical Engineering, Complex Systems and Biophysics. . The work reported in the thesis must represent a significant scientific advance. . If the thesis includes previously published material, permission to reproduce this must be gained from the respective copyright holder (a maximum 30% of the thesis should be a verbatim reproduction from the author’s previous publications). . They must have been examined and passed during the 12 months prior to nomination. . Each thesis should include a foreword by the supervisor outlining the significance of its content. . The theses should have a clearly defined structure including an introduction accessible to new PhD students and scientists not expert in the relevant field. Indexed by zbMATH.

Andre Filipe Furtado

Seismic Vulnerability Assessment and Retrofitting Strategies for Masonry Infilled Frame Building Doctoral Thesis accepted by University of Porto, Porto, Portugal

Author Dr. Andre Filipe Furtado Department of Civil Engineering and Architecture CERIS - Instituto Superior Técnico Lisboa, Portugal

Supervisor Prof. António Arêde Department of Civil Engineering Faculty of Engineering University of Porto Porto, Portugal Co-supervisor Prof. Hugo Rodrigues Department of Civil Engineering University of Aveiro Aveiro, Portugal

ISSN 2190-5053 ISSN 2190-5061 (electronic) Springer Theses ISBN 978-3-031-20371-8 ISBN 978-3-031-20372-5 (eBook) https://doi.org/10.1007/978-3-031-20372-5 © The Editor(s) (if applicable) and The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 This work is subject to copyright. All rights are solely and exclusively licensed by the Publisher, whether the whole or part of the material is concerned, specifically the rights of translation, reprinting, reuse of illustrations, recitation, broadcasting, reproduction on microfilms or in any other physical way, and transmission or information storage and retrieval, electronic adaptation, computer software, or by similar or dissimilar methodology now known or hereafter developed. The use of general descriptive names, registered names, trademarks, service marks, etc. in this publication does not imply, even in the absence of a specific statement, that such names are exempt from the relevant protective laws and regulations and therefore free for general use. The publisher, the authors, and the editors are safe to assume that the advice and information in this book are believed to be true and accurate at the date of publication. Neither the publisher nor the authors or the editors give a warranty, expressed or implied, with respect to the material contained herein or for any errors or omissions that may have been made. The publisher remains neutral with regard to jurisdictional claims in published maps and institutional affiliations. This Springer imprint is published by the registered company Springer Nature Switzerland AG The registered company address is: Gewerbestrasse 11, 6330 Cham, Switzerland

Supervisor’s Foreword

The Reinforced Concrete (RC) buildings comprise about 60% of the building stock and host approximately 65% of the population. Half of these buildings were not designed according to modern seismic codes. Recent earthquakes have evidenced the significant role of the masonry infill walls. In particular, a high vulnerability of the infill walls was observed under out-of-plane loadings. Based on this motivation the work developed within this thesis focused on an extensive experimental characterization of the masonry infill walls out-of-plane seismic behaviour. Three different experimental tests were performed: (i) material characterization tests; (ii) mechanical characterization tests and (iii) in-plane and out-of-plane tests on full-scale specimens. Different types of masonry infill walls were studied (i.e. other masonry units, support conditions and workmanship). The experimental work has significantly impacted the literature since it proved the infill walls’ high seismic vulnerability of infill walls when subjected to out-of-plane loadings. An innovative testing platform was developed to carry out out-of-plane tests of walls with different geometry types, with and without openings. In addition, a simplified macro-model was proposed to simulate the seismic behaviour of infill walls and their interaction with the envelope reinforced concrete structures. Several parametric studies were performed proving that neglecting the presence of the infill walls or even their outof-plane behaviour can result in catastrophic consequences for the building (i.e. collapse) in future earthquakes. Porto, Portugal April 2021

Prof. António Arêde

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Parts of this thesis have been published in the following documents Journals A. Furtado, H. Rodrigues and A. Arêde (2021) “Effect of the infill panels in the floor response spectra of an 8-storey RC building”, Structures, vol. 34, Pages 2476–2498, DOI: https://doi.org/10.1016/j.istruc.2021.08.102 (Scopus indexed) A. Furtado, H. Rodrigues, A. Arêde, J. Melo and H. Varum (2021) “The use of textile-reinforced mortar as a strengthening technique for the infill walls out-ofplane behaviour”, Composite Structures, Vol. 255, DOI: https://doi.org/10.1016/j. compstruct.2020.113029 (In Press, Scopus indexed) A. Furtado, H. Rodrigues and A. Arêde (2021) “Cantilever flexural strength tests of masonry infill walls strengthened with textile-reinforced mortar”, Journal of Building Engineering, Vol. 33, DOI: https://doi.org/10.1016/j.jobe.2020.101611 (Scopus indexed) A. Furtado, H. Rodrigues, A. Arêde and H. Varum (2020) “Experimental tests on strengthening strategies for masonry infill walls: A literature review”, Construction and Building Materials, Vol. 263, DOI: https://doi.org/10.1016/j.conbuildmat.2020. 120520 (Scopus indexed) A. Furtado, H. Rodrigues, A. Arêde, H. Varum (2020) “Mechanical properties characterization of different types of masonry infill walls” Frontiers of Structural and Civil Engineering, DOI: https://doi.org/10.1007/s11709-019-0602-y A. Furtado, H. Rodrigues, H. Varum and A. Arêde (2018) “Mainshock-aftershock damage assessment of infilled RC structures”, Engineering Structures, Volume 175, 2018, Pages 645–660, ISSN 0141-0296, DOI: https://doi.org/10.1016/j.engstruct. 2018.08.063 (Scopus indexed) A. Furtado, H. Rodrigues, A. Arêde and Humberto Varum “Modal identification of infill masonry walls with different characteristics”, Engineering Structures, Volume 145, Pages 188–134, DOI: https://doi.org/10.1016/j.engstruct.2017.05.003 (Scopus indexed) A. Furtado, H. Rodrigues, A. Arêde and H. Varum (2016) “Experimental evaluation of out-of-plane capacity of masonry infill walls”, Engineering Structures, Volume 111, 15 March 2016, Pages 48–63, ISSN 0141-0296, DOI: https://doi.org/10.1016/ j.engstruct.2015.12.013 (Scopus indexed)

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Parts of this thesis have been published in the following documents

International Conferences A. Furtado, H. Rodrigues, A. Arêde and H. Varum (2018) “Assessment of the Mainshock-Aftershock Collapse Vulnerability of RC Structures Considering the in-Plane and Out-of-Plane Behaviour”, 16th European Conference on Earthquake Engineering (16ECEE), June 18–21, Thessaloniki, Greece A. Furtado, H. Rodrigues, A. Arêde and H. Varum (2018) “Out-of-Plane Cyclic Performance of Full-Scale Infill Masonry Walls Subjected to Out-of-Plane Loadings Using Airbags”, 16th European Conference on Earthquake Engineering (16ECEE), June 18–21, Thessaloniki, Greece A. Furtado, H. Rodrigues, A. Arˆede and H. Varum (2017) “Characterization of the out-of-plane behaviour of full-scale masonry infill walls with and without previous in-plane damage: experimental study”, 16th World Conference on Earthquake Engineering, 16WCEE 2017, 9–13 January, Santiago, Chile

Acknowledgements

Since the day I first came to the Faculty of Engineering of the University of Porto, I felt that this was my home, that this was a school where I could grow up as a student, as an engineer but above all as a man. Proof of that is the fact that with this thesis, I am reaching the top of a mountain that sometimes was so hard to climb, but it was the most beautiful journey of my life until the moment. It has been a fruitful journey full of new academic and personal experiences. However, like any other experience in our life, it would not have been possible without the contribution and support of the kind people who accompanied me throughout this process. In the following paragraphs, I would like to express my most sincere gratitude and appreciation for those who have helped me complete this journey. First of all, I would like to express my gratitude to Prof. António Arêde. Acknowledge him for giving me the opportunity in 2013 to join in this long and challenging journey. I will never forget all the opportunities, confidence, and knowledge that He shared with me. For all the friendship, support and encouragement throughout this journey. For being an exceptional Professor and Supervisor. Thanks for sharing all the love for the experimental field and earthquake engineering with me. Second, I would like to express special gratitude to Prof. Hugo Rodrigues. More than my co-supervisor, one of my best friends during these years. I will never forget all the support every day during this period. All the advices, the concern, the friendship. I consider myself a lucky person for have a person like Hugo in my life. An inspiration and an example for me. He always knows how to push my best capacities and how to motivate me. For all the beers, all the late-night hours working, for all the adventures during all these years, and for other thousand reasons that I think it is not possible to express in words, thank you. To Prof. Humberto Varum, his strong and close support made it possible to complete this Ph.D. I will not forget all the efforts to ensure that everything was going well. A very special acknowledgement to Telma Cruz, my wife and love of my life, who has inspired me during all this journey. I would like to acknowledge all the love, friendship, patience and understanding. Thanks for helping me to be better day after day. I am sure that the rest of our lives will be magical. ix

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Acknowledgements

Because family is the support of our lives, this work would not be possible without the support and inspiration of my mother and my grandparents. All of them were present in each word of this thesis. They were the base of my force and ambition since I was born. Thanks for your love and for being always there for me. A large part of this Ph.D. was carried out at the Laboratory for Earthquake and Structural Engineering. During all this period, exceptional persons helped me develop all the experimental work, namely, Mr. Valdemar Luís, Mr. Guilherme Nogueira and Mr. Nuno Pinto. I would also like to express my gratitude to Preceram, Cimpor and Fassa Bortolo for supporting all the experimental campaigns presented by supplying all the material. The author is grateful for the Foundation for Science and Technology’s support through funding UIDB/04625/2020 from the research unit CERIS.

Contents

1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.1 Framework and Motivation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.2 Objectives and Tasks . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.3 Thesis Organization . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2 Out−of−Plane Behaviour of Masonry Infill Walls Experimental Tests: A Systematic Review . . . . . . . . . . . . . . . . . . . . . . . . 2.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.2 Methodology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.3 Selection Criteria . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.4 Study Selection . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.5 Data Extraction, Quality Assessment and Data Synthesis . . . . . . . . 2.6 Final Database . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.7 Global Overview of OOP Tests on Infill Masonry Walls . . . . . . . . . 2.8 OOP Load Application Strategies and Protocol . . . . . . . . . . . . . . . . 2.9 OOP Behaviour of As-Built Infill Masonry Walls . . . . . . . . . . . . . . 2.9.1 Effect of RC Frame Type and Detailing . . . . . . . . . . . . . . . 2.9.2 Effect of Geometric Dimensions and Mechanical Properties . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.9.3 Effect of Gravity Load . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.9.4 Effect of Openings and Panel Support Conditions . . . . . . . 2.10 Assessment of the Previous IP Damage Effect on the Infill Walls OOP Capacity . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.10.1 Background . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.10.2 Cracking Strength and Secant Cracking Stiffness . . . . . . . 2.10.3 Maximum Strength and Secant Stiffness . . . . . . . . . . . . . . 2.10.4 IP−OOP Ultimate Displacement . . . . . . . . . . . . . . . . . . . . . 2.11 Final Considerations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

1 1 3 5 7 9 9 10 10 10 11 13 13 17 22 22 23 27 28 29 29 31 34 38 39 40

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3 Strengthening Strategies to Improve the Seismic Behaviour of Infill Masonry Walls . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.2 Literature Review of Retrofit and Strengthening Techniques . . . . . 3.2.1 Disconnection of the Infills from the Structural System . . . . . . . . . . . . . . . . . . . . . . . . . . 3.2.2 Infill Masonry Walls Retrofitting and Strengthening Techniques . . . . . . . . . . . . . . . . . . . . . . . 3.2.3 Reinforced Plaster . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.2.4 Bed Joints Reinforcement . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.3 Final Considerations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4 Experimental Characterization of the As-Built Masonry Infill Components’ Properties . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.1 Framework . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.2 Stage 1—Material Characterization Tests of Masonry Units and Mortar . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.2.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.2.2 Compression Strength Tests of Masonry Units . . . . . . . . . 4.2.3 Compression and Flexural Strength Tests of the Mortar . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.3 Stage 2—Mechanical Characterization of the Infill Masonry Wallets . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.3.1 Testing Campaign Overview . . . . . . . . . . . . . . . . . . . . . . . . . 4.3.2 Mechanical Properties of Mortar . . . . . . . . . . . . . . . . . . . . . 4.3.3 Compressive Strength Tests . . . . . . . . . . . . . . . . . . . . . . . . . 4.3.4 Diagonal Tensile Strength Tests . . . . . . . . . . . . . . . . . . . . . . 4.3.5 Flexural Strength Tests . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.3.6 Final Remarks on Stage 2 Tests . . . . . . . . . . . . . . . . . . . . . . 4.4 Stage 3—Dynamic Characterization of Infill Panels Through Ambient Vibration Tests . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.4.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.4.2 Methodology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.4.3 Laboratory Tests . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.4.4 In-Situ Tests . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.4.5 Final Remarks on Stage 3 Tests . . . . . . . . . . . . . . . . . . . . . . 4.5 Final Considerations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

45 45 45 46 50 61 74 78 79 83 83 88 88 89 94 97 97 98 99 108 113 122 124 124 125 127 132 153 155 156

5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built Masonry Infill Walls . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 159 5.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 159 5.2 Specimens Details . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 160

Contents

5.3

Material Properties and Specimen Construction . . . . . . . . . . . . . . . . 5.3.1 Concrete and Reinforcement Steel Bars . . . . . . . . . . . . . . . 5.3.2 Mortar and Masonry Units . . . . . . . . . . . . . . . . . . . . . . . . . . 5.4 Stage 1—Out-of-Plane Tests Using Airbags . . . . . . . . . . . . . . . . . . . 5.4.1 Test Setup . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.4.2 Instrumentation and Loading Protocol . . . . . . . . . . . . . . . . 5.4.3 Stage 1: Individual Test Results . . . . . . . . . . . . . . . . . . . . . . 5.4.4 Stage 1: Global Results’ Comparison . . . . . . . . . . . . . . . . . 5.5 Stage 2—Out-of-Plane Tests Using Pneumatic Actuators . . . . . . . . 5.5.1 Test Setup . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.5.2 Instrumentation and Loading Protocol . . . . . . . . . . . . . . . . 5.5.3 Stage 2: Individual Test Results . . . . . . . . . . . . . . . . . . . . . . 5.5.4 Stage 2: Global Results Comparison . . . . . . . . . . . . . . . . . . 5.6 Overview of the Summary of Global Results: Critical Analysis . . . 5.6.1 Effect of Columns Axial Load . . . . . . . . . . . . . . . . . . . . . . . 5.6.2 Effect of Reduction of the Panel Support Width . . . . . . . . 5.6.3 Effect of Previous Damage . . . . . . . . . . . . . . . . . . . . . . . . . . 5.6.4 Effect of Plaster . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.6.5 Effect of Workmanship . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.6.6 Effect of Test Setup . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.7 Final Remarks and Major Observations from the Experimental Campaign . . . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6 Experimental Assessment of Strengthening Solutions to Prevent the OOP Collapse of Infill Masonry Walls Through Textile Reinforced Mortars . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.2 Flexural Strength Tests in Strengthened Infill Masonry Wallets . . . 6.2.1 Objectives and Specimens’ Description . . . . . . . . . . . . . . . 6.2.2 Specimens Description and Strengthening Process . . . . . . 6.2.3 Test Setup Description and Instrumentation . . . . . . . . . . . . 6.2.4 Experimental Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.3 Experimental Assessment of TRM Solutions to Improve the Out-of-Plane Capacity of Full-Scale Infill Masonry Walls . . . . 6.3.1 Objectives and Specimens’ Description . . . . . . . . . . . . . . . 6.3.2 Description of the Strengthening Strategies . . . . . . . . . . . . 6.3.3 Material Properties . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.3.4 Test Setup, Instrumentation and Loading Protocol . . . . . . 6.3.5 Experimental Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.3.6 Comparison of the Results and Discussion . . . . . . . . . . . . . 6.4 Final Considerations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

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163 163 164 167 167 169 170 178 193 193 193 197 201 210 213 214 215 218 219 220 220 222

223 223 224 224 225 227 231 234 234 236 242 243 243 251 260 262

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7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.2 Brief Overview on Numerical Modelling Approaches: From Simplified Macro-models to Detailed Modelling . . . . . . . . . . . . . . . 7.2.1 Simplified Macro-models . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.2.2 Detailed Micro-modelling Approach . . . . . . . . . . . . . . . . . . 7.3 Development of a Simplified Modelling Approach to Simulate Infill Walls in OpenSees . . . . . . . . . . . . . . . . . . . . . . . . . 7.3.1 General Considerations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.3.2 Modelling of the IP Behaviour . . . . . . . . . . . . . . . . . . . . . . . 7.3.3 Modelling of the OOP Behaviour . . . . . . . . . . . . . . . . . . . . 7.4 Numerical Modelling Calibration . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.4.1 In-Plane Behaviour—Panels Without Openings . . . . . . . . 7.4.2 In-Plane Behaviour—Panels with Openings . . . . . . . . . . . . 7.4.3 OOP Behaviour . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.5 Case Study . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.5.1 Description of the Building . . . . . . . . . . . . . . . . . . . . . . . . . . 7.5.2 Numerical Modelling Strategy . . . . . . . . . . . . . . . . . . . . . . . 7.5.3 Material Modelling Properties . . . . . . . . . . . . . . . . . . . . . . . 7.5.4 Preliminary Results, Natural Frequencies and Vibration Modes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.5.5 Effect of the Infill Walls in the Seismic Vulnerability Assessment of Rc Buildings Parametric Study 1 . . . . . . . . 7.5.6 Effect of the Infill Panels in the Floor Response Spectra Parametric Study 2 . . . . . . . . . . . . . . . . . . . . . . . . . . 7.5.7 Seismic Vulnerability Assessment of Infilled RC Structures Subjected to Mainshock-Aftershock Sequence Parametric Study 3 . . . . . . . . . . . . . . . . . . . . . . . . 7.6 Final Considerations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

263 263 264 265 268 269 269 271 274 276 276 284 291 293 293 295 295 297 298 307

327 339 341

8 Final Conclusions and Future Works . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 347 8.1 Conclusions and Major Contributions . . . . . . . . . . . . . . . . . . . . . . . . 347 8.2 Future Works . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 350 Author Biography . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 353 Bibliography . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 355

Chapter 1

Introduction

1.1 Framework and Motivation The presence of infill masonry (IM) walls in reinforced concrete (RC) buildings is very common. The infill walls are widely used for partition purposes and provide thermal and acoustic insulation to the buildings. However, and even today, in the design process of new buildings and the structural safety assessment of existing ones, the infills are usually considered as “non-structural” elements, and their influence on the structural response is disregarded. The infill walls’ presence is recognized to have an important role in the global structural behaviour and the performance of RC frame building structures when subjected to earthquake demands. This has been proved in recent earthquakes all over the world, where many RC buildings observed poor seismic performance with extensive damage to the structural and non-structural elements (see Fig. 1.1). Furthermore, the last events showed that a crucial issue for life-safety and loss reduction due to earthquakes in existing RC buildings is related to mitigating the out-of-plane (OOP) collapse of infill masonry walls. The large shear demands that infill walls may attract and the corresponding inplane (IP) damage evolution are likely to increase their OOP vulnerability. The eventual collapse of infill walls OOP, see Fig. 1.2, can result in human severe injuries and casualties and high economic losses. Therefore, adequate knowledge of all the aspects related to infilled framed structures’ behaviour, their components (structural and non-structural elements) and the phenomena interaction is fundamental to guide the designers in assessing and strengthening existing buildings. Due to their interacion with the surrounding RC frame, infill walls can develop enhanced OOP strength through the arching mechanism, which mainly depends on the panel slenderness, masonry compressive strength, boundary conditions and panel support width conditions. The seismic vulnerability of the infill walls can increase due to constructive detailing aspects commonly adopted throughout Southern European countries. For example, the following issues can contribute to the reduction of the infill walls OOP © The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 A. F. Furtado, Seismic Vulnerability Assessment and Retrofitting Strategies for Masonry Infilled Frame Building, Springer Theses, https://doi.org/10.1007/978-3-031-20372-5_1

1

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1 Introduction

Fig. 1.1 Infilled RC structure damages after L’Aquila earthquake in 2009 [1]

a)

b)

Fig. 1.2 Infill panels OOP collapse after: a Lorca earthquake in 2009 [9] and b L’Aquila earthquake in 2009 [1]

strength and deformation capacity and thus, increase the risk of collapse: (i) no connection between the panel and the surrounding RC elements; (ii) no connection between the internal and external leafs (in the case of double-leaf infill walls) and, (iii) reduced width of the panel support. In this framework, experimental studies are important to understand the OOP behaviour of the infill panels, especially for OOP failure combined with prior IP damages, and numerical models are fundamental to studying their effect on the global building behaviour. The investigation of this issue deepened in the 90s [2], concluding that the OOP strength reduces with the panel slenderness. Some authors also observed that the

1.2 Objectives and Tasks

3

panel aspect ratio could significantly affect the OOP behaviour [3]. The IP-OOP behaviour interaction seems to be one of the most important issues, since different authors report that the previous damage reduces the panel OOP strength [4, 5]. The boundary conditions also affects the panel OOP behaviour, since it modify the failure mechanism and their OOP strength [6]. The workmanship’s role is also pointed out as a variable that could introduce variability regarding the panel OOP behaviour response [7, 8]. However, taking into account the lack of experimental studies carried out over the last years regarding the OOP behaviour of infill walls and considering, in particular, the common use of hollow clay horizontal bricks in Portugal, as well as in other European Countries, in the last decades, it is of utmost importance to evaluate its seismic performance and to propose strengthening strategies. For assessing RC frame structures, the nonlinear behaviour induced by earthquake demands and the influence of infill walls should be considered [10–19], for which different numerical modelling strategies to simulate the infill panels’ seismic response can be found in the literature [20–22], ranging from refined micro-models to simplified macro-models. The micro-models involve a high discretization level of the panel, while the macro-models rely on several simplifications aiming to represent the panel’s global behaviour with a few structural elements and mechanical parameters. In many situations, it is unsuitable to adopt refined models for non-linear analysis of complex structures under earthquake action. Thus, for simulating the response of frame structures with infill walls and considering the interaction between them, the adoption of simplified models is unavoidable [23, 24]. This is further confirmed in [25], which recommends assessing the structural response of buildings considering the infill panels represented by an equivalent diagonal strut model.

1.2 Objectives and Tasks Considering the previously presented background, the main goals of the present thesis consist of the following: . Experimental characterization of the infill panels material, mechanical and dynamic properties; . Experimental assessment of the OOP behaviour of full-scale infill panels with and without previous damage; . Evaluation of the efficiency of strengthening solutions to improve their seismic behaviour and prevent the OOP collapse; . Development and implementation of a simplified macro-modelling strategy to simulate the panels’ seismic behaviour and their interaction with the surrounding frame; . Assessment of the effect of the infill panels in the seismic response of RC structures. In order to achieve these objectives, the experimental, analytical and numerical work subjects of this thesis are organized into five major tasks, as shown in Fig. 1.3.

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1 Introduction

Fig. 1.3 Schematic layout of the work plan framework: sequence and interaction between tasks

The first task is related to the state-of-art review, which is sub-divided into four subtasks. First, information collected from post-earthquake survey assessment reports is presented and discussed. A compilation of the most common typologies of damages focuses on the infill walls’ effect in RC buildings’ seismic response. The economic losses due to the infills in recent earthquakes are also discussed. The second subtask is related to the review of experimental tests of infill panels subjected to OOP loadings carried out in the last years, following a systematic review approach. The effect of different variables in the panel response is analysed (in support of task 2). The third sub-task addresses the literature review of strengthening techniques of infill panels (to support task 3). The fourth and last sub-task reviews existing numerical modelling strategies to simulate the infill panels’ seismic behaviour (in support of tasks 4 and 5). Task 1 is planned to be the basis that supports the development of all the other tasks, since it provides the review of the knowledge available in the literature that is necessary for the development of this thesis. Task 2 is related to the experimental characterization of as-built masonry infill walls, which is divided into four sub-tasks, namely: (i) material characterization of masonry units and mortars; (ii) mechanical characterization of infill components which comprises compression strength tests, diagonal tensile strength tests and flexural strength tests, parallel and perpendicular to the horizontal bed joints; these properties provide support for the modelling and calibration of the infill panels in task 4 and 5; (iii) characterization of the dynamic properties of infill panels resorting to ambient vibration tests; the results herein obtained allow calibrating the numerical models in Task 4; and (iv) experimental characterization of the OOP behaviour of full-scale infill panels for assessing the influence of variables such as prior damage due to previous IP tests, reduction of the panel width support, columns axial load, plaster, workmanship and test setup; as a key issue, this sub-task comprises the development of the test setup for the execution of the OOP tests.

1.3 Thesis Organization

5

Task 3 is dedicated to the study, definition and testing of strengthening strategies for infill walls. The task is divided into two sub-tasks: (i) experimental campaign of flexural strength tests in strengthened infill masonry wallets; and (ii) experimental campaign of OOP tests in strengthened full-scale infill walls with and without prior damage due to IP tests. Task 4 covers developing and implementing a simplified macro-model to simulate the infill walls’ seismic behaviour in the software OpenSees [26]. Therefore, this task includes the calibration of the numerical modelling approach and provides support to Task 5. Finally, task 5 is related to parametric studies to assess the impact of the infill panels in the seismic response of RC structures. For that purpose, this task is divided into three following sub-tasks: (i) influence of the infills in the seismic assessment of RC buildings; (ii) effect of the infills in the floor seismic response spectra; and (iii) study of the seismic vulnerability of infilled RC structures subjected to the mainshock-aftershock sequence.

1.3 Thesis Organization The present thesis is divided into nine chapters complemented with the list of references introduced throughout the text. In brief, Chaps. 2 and 3 consist of state-of-art reviews concerning the existing OOP tests of infill panels and the strengthening strategies for infill panels, respectively; Chapter 4 presents an experimental characterization of the material, mechanical and dynamic properties of infill panels built with hollow clay horizontal bricks while Chap. 5 describes the experimental characterization of as-built infill panels OOP behaviour and Chap. 6 focuses on the experimental assessment of textilereinforced mortar (TRM) strengthening solutions to improve the infill panels’ seismic behaviour; Chapter 7 presents the development, implementation and calibration of a simplified numerical model to simulate the infill panels’ seismic behaviour and, finally, Chapter 8 draws the major contributions and conclusions of the work developed in this thesis and indicates future research works. A more detailed description of the overall contents of each chapter is proved next. In the present chapter the background of the current work is introduced, and the main goals are described. The structure of this thesis is also described. Chapter 2 addresses a systematic review of experimental tests available in the literature that were performed to study the infill walls OOP behaviour. Open challenges still lacking deeper research and discussion are pointed out. A systematic review methodology is presented and detailed. The final list of research works that were selected to be studied within the framework of this thesis is presented. Chapter 3 presents an extensive review of infill masonry strengthening strategies available in the literature. The strategies are divided according to two different approaches, namely: (a) disconnection of the panel from the surrounding RC frame

6

1 Introduction

and, (b) panel strengthening. The design details of each solution, the strengthening material properties and the respective experimental results and major findings are presented and discussed. Chapter 4 presents a large experimental campaign comprising material and mechanical characterization tests as well as dynamic characterization tests that were carried out. The material characterization tests were performed in the masonry units used for this thesis works. Mechanical characterization was divided into four types of tests: (i) compressive strength tests; (ii) diagonal tensile strength tests; (iii) flexural strength tests parallel to the horizontal bed joints; and (iv) flexural strength tests perpendicular to the horizontal bed joints. Concerning the dynamic characterization, ambient vibration tests were carried out under laboratory conditions and in-situ. The testing details, major results and conclusions are provided throughout this chapter. Chapter 5 presents an experimental campaign of full-scale infill walls OOP tested quasi-statically with airbags and later with pneumatic jacks. A set of nine specimens were tested to assess the influence of different parameters such as: type of OOP loading (monotonic or half-cyclic); axial load on columns; previous damage due to IP test; panel support width; presence of plaster; workmanship and test setup. Details regarding the test setup, instrumentation, loading protocol, major experimental results and findings are also provided in this chapter. Chapter 6 describes the experimental assessment of the efficiency of TRM based solutions to improve the infill panels’ OOP behaviour. First, a testing campaign of flexural strength tests was carried out in infill wallets strengthened with two different solutions. The major goal was to evaluate the effect of the strengthening mesh and connectors. The second part of the chapter is related to the OOP experimental testing of full-scale strengthened panels. Again, two different strengthening strategies are tested in which it is evaluated the effect of the mesh and its anchorage to the RC frame. The campaign includes the testing of three strengthened specimens, with and without prior damage due to IP test, which results are compared with those obtained in tests of as-built specimens presented in Chap. 6. In Chap. 7, after a brief review of numerical modelling approaches available to simulate the infill panels’ seismic behaviour, it is presented a simplified macromodel developed within the present work in order to simulate the infill walls’ seismic behaviour. The simplified macro-model allows the simulation of the combined IPOOP non-linear behaviour and was implemented in the OpenSees software [26]. The major details concerning the numerical modelling, the adopted calibration methodology and corresponding implementation procedures in the OpenSees software [26] are described in this chapter. Furthermore, a case study is presented for which the seismic vulnerability assessment of an infilled RC structure is carried out and the impact of considering different infill walls’ modelling strategies is analysed. A parametric study was developed and presented, concerning the effect of the infill walls in the floor seismic response spectra of the structure. The characterization of the seismic behaviour of the infill panels during an earthquake event is also analysed and, finally, another parametric study is presented focusing on the seismic assessment of infilled RC structures subjected to mainshock-aftershock sequences.

References

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Chapter 8 closes this document with the summary of the most relevant outcomes and results achieved during this thesis work and the key results are identified concerning the seismic vulnerability assessment and strengthening of infill masonry panels. Finally, possible future works are also suggested.

References 1. Vicente R, Rodrigues H, Varum H, Costa A, Mendes da Silva R (2012) Performance of masonry enclosure walls: lessons learned from recent earthquakes. Earthq Eng Eng Vib 11(1):23–34 2. Dawe J, Seah C (1989) Out-of-plane resistance of concrete masonry infilled panels. Can J Civil Eng 16(6):854–864 3. De Risi MT, Di Domenico M, Ricci P, Verderame GM, Manfredi G (2019) Experimental investigation on the influence of the aspect ratio on the in-plane/out-of-plane interaction for masonry infills in RC frames. Eng Struct 189:523–540 4. Ricci P, Di Domenico M, Verderame GM (2018) Empirical-based out-of-plane URM infill wall model accounting for the interaction with in-plane demand. Earthq Eng Struct Dyn 47(3):802– 827 5. Ricci P, Di Domenico M, Verderame GM (2018) Experimental investigation of the influence of slenderness ratio and of the in-plane/out-of-plane interaction on the out-of-plane strength of URM infill walls. Constr Build Mater 191:507–522 6. Di Domenico M, Ricci P. Verderame GM (2018). Experimental assessment of the influence of boundary conditions on the out-of-plane response of unreinforced masonry infill walls. J Earthq Eng 1–39 7. Akhoundi F, Vasconcelos G, Lourenço P (2018) Experimental out-of-plane behavior of brick masonry infilled frames. Int J Arch Heritage 14(2):221–237 8. Akhoundi F, Vasconcelos G, Lourenço P, Silva LM, Cunha F, Fangueiro R (2018) In-plane behavior of cavity masonry infills and strengthening with textile reinforced mortar. Eng Struct 156:145–160 9. Romão X, Costa AA, Paupério E, Rodrigues H, Vicente R, Varum H, Costa A (2013) Field observations and interpretation of the structural performance of constructions after the 11 May 2011 Lorca earthquake. Eng Fail Anal 34:670–692 10. Hak S, Morandi P, Magenes G, Sulivan TJ (2012) Damage control for clay masonry infills in the design of RC frame structures. J Earthq Eng 16(1):1–35 11. Furtado A, Rodrigues H, Arêde A (2015) Modelling of masonry infill walls participation in the seismic behaviour of RC buildings using OpenSees. Int J Adv Struct Eng (IJASE) 7(2):117–127 12. Akhoundi F, Lourenço P, Vasconcelos G (2016) Numerically based proposals for the stiffness and strength of masonry infills with openings in reinforced concrete frames. Earthq Eng Struct Dyn 45(6):869–891 13. Akhoundi F, Vasconcelos G, Lourenço P, Silva B (2016) Out-of-plane response of masonry infilled RC frames: effect of workamnship and opening. International Brick and Block Masonry Conference—IB2MAC. Padua, Italy 14. Furtado A, Costa C, Arêde A, Rodrigues H (2016) Geometric characterisation of Portuguese RC buildings with masonry infill walls. Eur J Environ Civil Eng 1–16 15. Furtado A, Rodrigues H, Arêde A, Varum H (2016) Experimental evaluation of out-of-plane capacity of masonry infill walls. Eng Struct 111:48–63 16. Furtado A, Rodrigues H, Arêde A, Varum H (2016) Simplified macro-model for infill masonry walls considering the out-of-plane behaviour. Earthq Eng Struct Dyn 45(4):507–524 17. Fardis M, Panagiotakos T (1997) Seismic design and response of bare and masonry-infilled reinforced concrete buildings: part II: Infilled structures. J Earthq Eng 13:475–503 18. Dolsek M, Fajfar P (2005) Simplified non-linear seismic analysis of infilled reinforced concrete frames. Earthq Eng Struct Dyn 34:49–66

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19. Dolsek M, Fajfar P (2008) The effect of masonry infills on seismic response of a four storey reinforced concrete frame - a probabilistic assessment. Eng Struct 30:1991–2001 20. Asteris P, Antoniou S, Sophianopoulos D, Chrysostomou C (2011) Mathematical macromodeling of infilled frames: state of the art. J Struct Eng 137(12):1508–1517 21. Koutromanos I, Stavridis A, Shing P, Willam K (2011) Numerical modeling of masonry-infilled RC frames subjected to seismic loads. Comput Struct 89:1026–1037 22. Asteris P, Cotsovos D, Chrysostomou C, Mohebkhah A, Al-Chaar G (2013) Mathematical micromodelling of infilled frames: State of the art. Eng Struct 56:1905–1921 23. Rodrigues H, Varum H, Costa A (2010) Simplified macro-model for infill masonry panels. J Earthq Eng 14(3):390–416 24. Smyrou E, Blandon C, Antoniou S, Pinho R, Crisafulli F (2011) Implementation and verification of a masonry panel model for nonlinear dynamic analysis of infilled RC frames. Bull Earthq Eng 9:1519–1534 25. FEMA-356 (2000) Pre-standard and commentary for the seismic rehabilitation of buildings F. E. M. Agency. Washington, DC 26. Mckenna F, Fenves G, Scott M, Jeremic B (2000) Open system for earthquake engineering simulation (OpenSees). Berkley, CA

Chapter 2

Out−of−Plane Behaviour of Masonry Infill Walls Experimental Tests: A Systematic Review

2.1 Introduction Over the literature, there is a small number of testing campaigns related to the study and characterization of the OOP behaviour of infill panels of steel or RC frames [1–14]. Part of these testing campaigns consisted on shaking table tests of single infill panels or scaled infilled RC structures [15–22]. The main scope of this chapter is to present a systematic review of the experimental tests available in the literature that were performed to study the OOP behaviour of infill walls. A systematic review methodology will be detailed and the final group list of works that will be studied within the framework of this thesis will be presented in the form of a database. Similar databases were developed by other authors in the field of mechanical modelling of existing masonry assemblages and earthquake performance of infilled frames, among others [23, 24]. With this systematic review, the following information will be collected from each specimen: panel geometry, slenderness, aspect ratio, openings, boundary conditions, masonry compressive and flexural strength, masonry units and respective percentage of voids, mortar compressive and flexural strength, loading protocol (gravity load and type of OOP loading) and existence of previous damage due to previous IP test. Discussion regarding the effects of each parameter and the major findings obtained by the different authors will be addressed throughout the chapter. Finally, it will be also pointed out the open challenges that are still requiring deeper research and discussion.

© The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 A. F. Furtado, Seismic Vulnerability Assessment and Retrofitting Strategies for Masonry Infilled Frame Building, Springer Theses, https://doi.org/10.1007/978-3-031-20372-5_2

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2 Out−of−Plane Behaviour of Masonry Infill Walls Experimental Tests: …

2.2 Methodology 2.3 Selection Criteria Throughout the literature, different types of experimental testing are reported having been carried out to characterize the infill panels’ OOP behaviour. From quasi−static monotonic/cyclic tests until shaking table tests on infilled frames, the main goal was to characterize the seismic performance of these elements when subjected to pure OOP loadings or combined IP−OOP ones. Inspired in the works developed by Higgins et al. [25, 26] in the field of medical studies, a systematic review methodology was used to filter the most relevant and complete works available in the literature concerning the infill masonry walls OOP tests. The systematic review methodology can refine the literature by using explicit, rigorous, and reproducible methods to identify, critically appraise and synthesize the best evidence from a specific subject. This methodology allows refining the source of specific topics and better assessing the influence of certain variables in the scope of a research work, which can be useful in the structural engineering research field to analyse the available experimental data, mainly in specific topics with reduced amount of data. Based on that, this methodology started with a comprehensive literature search that was conducted on four electronic indexing databases: Web of Knowledge, Science Direct, Civil Engineering database (ASCE) and Scopus. In this literature review, the most up to date research works, contemporary with this thesis, were collected. Some studies available in the literature were also posterior to the tests carried out and published within this thesis (detailed in Chap. 5). Due to this, the experimental results obtained by the author were also used for this systematic review. Those results will be properly presented, detailed and discussed in Chap. 5. The following terms (keywords) were considered into the databases consulted: masonry infill walls; masonry infill test; out−of−plane masonry infill; cyclic tests masonry infill; shake table tests masonry infill and masonry infill out−of−plane cyclic behaviour.

2.4 Study Selection After retrieved, the records were combined using End Note (version Desktop X8) [27] and the duplicate ones were removed. At this point, titles and abstracts of the remaining records were screened to find relevant papers and their subsequent selection was made based on a set of criteria. A time window was set and only studies published between 1960 and 2019 were selected. Papers whose titles were not focused on infill masonry experimental tests were excluded. The full texts relative to the relevant titles/abstracts were read carefully and those who matched with the established eligibility criteria were included in the final database of papers. Numerical

2.5 Data Extraction, Quality Assessment and Data Synthesis

11

modelling research works were excluded from the final database. The testing of RC building structures (shake−table and pseudo−dynamic tests) were also excluded. Additionally, the most relevant document records such as PhD thesis and technical research reports were collected due to their high citation score and work novelty at the time that they were developed. In fact, it was considered appropriate to include these works due to the highly valuable and numerous information concerning this topic. A complete database search combined with other sources of record retrieval resulted in the identification of 990 documents. Following the elimination of duplicates, 533 papers remained to be screened. From these, only 97 were recognized as relevant for this topic, as illustrated in Fig. 2.1. Some papers were excluded in this phase, due to: publication form, i.e., technical notes or similar reports (n = 7); insufficient information of the test campaign (n = 16); tests of unreinforced masonry (URM) walls (n = 3), studies of combined experimental and numerical simulation of OOP behaviour of the infill masonry walls (n = 3).

2.5 Data Extraction, Quality Assessment and Data Synthesis Aiming at understanding the reasons behind the results variability in the experimental tests carried out in the infill masonry walls, several parameters and information were extracted from the final set of works, namely: • Generic information: geometric characteristics (panel height–Hinf , width–Winf and thickness t), number of tests, masonry unit, type of test, type of frame, frame details (design and drawings), test setup, OOP loading condition, gravity load; previous damage due to IP test and panel mass; • Material properties: type of mortar; mortar compressive (fmo ) and flexural strength (ft,mo ), percentage of voids of masonry unit (% of voids), compressive strength of masonry units (bricks) parallel (fb,parallel,unit ) and perpendicular (fb,perpendicular,unit ) to the holes direction; • Mechanical characterization: infill masonry compressive strength (fm,inf ), infill masonry elastic modulus (Em,inf ), infill masonry flexural strength parallel (fm,parallel,inf ) and perpendicular (fm,perpendicular,inf ) to the horizontal bed joints, infill masonry bond−shear strength and infill masonry diagonal tensile strength (Ss,inf ); Test results: the information extracted from the test results is illustrated in Fig. 2.2 and basically consists in the collection of the following parameters from the force– displacement curve: • Pre−Peak phase: point corresponding to 15% of the maximum strength before being reached the maximum peak load (F0.15Fmax , d0.15Fmax –triangle) • Pre−Peak phase: point corresponding to 30% of the maximum strength before being reached the maximum peak load (F0.30Fmax , d0.30Fmax –circle)

12

2 Out−of−Plane Behaviour of Masonry Infill Walls Experimental Tests: …

Records identified through

Additional Records identified through

database searching

other sources

(n=934)

(n=60)

Records after duplicates elimination (n=537)

Full-text papers excluded (n=7)

Insufficient information of the test Records screened

Records excluded

campaign

(n=101)

(n=52)

(n=16)

Study of unreinforced masonry walls OOP behaviour (not

Full-text papers assessed for eligibility (n = 53)

surrounded by envelope frames) (n=3)

Combined experimental and

Studies included in

numerical studies of IM walls OOP

qualitative synthesis

behaviour

(n=24)

(n=3)

Fig. 2.1 Systematic review methodology: selection of studies Fig. 2.2 Data information gathered from each specimen’ force–displacement result

2.7 Global Overview of OOP Tests on Infill Masonry Walls

13

• Peak phase: point corresponding to the maximum peak load (Fmax , dFmax –pentagram) and respective secant stiffness (Ksec –red dashed line which picks the pentagram point); • Post−Peak phase: point corresponding to 80% of the maximum strength after being reached the maximum peak load (F0.8Fmax , d0.8Fmax –pentagon) and respective secant stiffness (Ksec,0.8Fmax −blue dashed line which picks the pentagon point); • Post−Peak phase: point corresponding to the ultimate strength reached by the specimen, as defined by the authors (Fult , dult –diamond) and respective secant stiffness (Ksec,Fult −black dashed line which picks the diamond point); • Cracking point: point corresponding to the appearance of the first major cracking during the test−information provided by the authors (Fcrack , dcrack ) and respective secant stiffness (Ksec,crack ); • Cumulative energy dissipation and damages’ sequence observed during the test. The data extraction was independently analysed, and it was defined whether the data fulfilled or not the accuracy requirements. Evidence from the selected studies was summarized and hypotheses/explanations were derived. A qualitative analysis was adopted, since it was considered the most appropriate method to address the question of interest, while taking into account the heterogeneity of studies concerning the large number of parameters involved.

2.6 Final Database This systematic review methodology, encompassing the database search combined with other sources of record retrieval, resulted in the identification of 24 research works, which are summarized in Table 2.1. Additionally, aiming at assessing the amount of information provided by each work included in the database, a classification system was performed. The appraisal of classification items followed a dichotomous response format (present/absent), where the presence of each criterion was awarded one point and the absence was given zero points. Thus, a maximum score of 11 points can be reached with this methodology.

2.7 Global Overview of OOP Tests on Infill Masonry Walls From the tests available in the database, it can be observed that different types of tests were carried out to study the infill panels’ OOP behaviour, namely: quasi−static tests (monotonic or cyclic), pseudo−dynamic tests and shake−table tests. Since the main aim of this systematic review is to study of the panel behaviour only one−frame, one−storey specimens were considered for this analysis, as described

Liu [30]

Lunn et al. [50]

Varela−Rivera [31]

Pereira et al. [8]

Guidi et al. [32]

Hak [10]

7

8

9

10

11

12

5

6

7

6

14

2

4

4

Tu et al. [29]

Komaraneni et al. [20]

5

6

13

9

Angel et al [4]

Calvi et al. [7]

3

4

15

Dawe et al. [2, 3] 8

Frederiksen [28]

1

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

−

−

+

−

−

−

+

+

+

+

+

+

+/+/±

+/±/+

−/−/−/−

−/−/ ±

−/−/ ±

±/±

±/±

±/+/+

−/−/ ±

−/−/ ±

± /−

−/−/−/−

−/−/− −/−/−/−

+/+ /+

± /−

−/ ±

−/±

+/±

−/±

−/ ±

+/±

±/−

±/−

Number Specimens’ Test Loading Frame Material properties of tests details setup Protocol detailing Mortar Masonry (type, units (% of fm,mo ; voids; fm,mo ) specific weight; fm,parallel,unit ; funit,b,perp )

2

Author and year

+ / ± /−/−

+ / ± /−/−

+/+/+/±

± /−/−/−

−/−/−/−/−

−/−/−/−/−

+ / ± /−/ +

± /−/−/−

+ / ± /−/ +

+ / ± /−/−

±/+/±

+/±/±

+/+

+/+

+/+

+/+

+/+

−/ +

+/+

+/+

+/+

−/ +

−/ +

−/ +

−/−

−/−

−/−

−/−

−/−

−/−

−/−

−/−

±

−/−

−/−

−/−

−

−

−

−

−

−

−

−

−

−

−

−

+

+

+

−

+

−

+

+

−

−

−

+

5.7

5.4

8.6

4.5

5.6

4.1

7.7

7

7.2

5.8

5.7

7.2

Total

(continued)

Masonry Force–displacement K Energy Damages wallets (fm,inf ; (envelopes, global (Kini , dissipation observed results) Em,inf ; Ksec ) fb,paralell,inf , fb,perpendicular,inf , Ss,inf )

Test results

Table 2.1 Systematic review of infill walls OOP tests available in the literature: Final database and respective classification of each work selected

14 2 Out−of−Plane Behaviour of Masonry Infill Walls Experimental Tests: …

Moreno−Herrera 8 [36]

8

4

(Akhoundi et al. [37, 38]

Furtado [12]

Silva [39]

Di Domenico [40]

Akhoundi [41]

16

17

18

19

20

21

2

6

7

3

8

Preti [35]

Porto [11]

14

15

Singhal et al. [33, 4 34]

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

+

−

+

+

+

+/+ /+

+/±

+/+ /+

+/+ /+

± /−

+/±

+/+ /+

+/±

± /−

+/+/±

−/−/ + / +

−/−/ ±

+/+/±

+/+/±

−/−/ ±

−/−/−/−

±/+/+

± /−/−

Number Specimens’ Test Loading Frame Material properties of tests details setup Protocol detailing Mortar Masonry (type, units (% of fm,mo ; voids; fm,mo ) specific weight; fm,parallel,unit ; funit,b,perp )

13

Author and year

Table 2.1 (continued) Test results

+/+/+/+/ +

+ / ± /−/−

+/+/+/+/ +

+/+/+/+/ +

−/−/−/−/−

± /−/−/−

+ / + / ± /−

−/−/−/−

+/+/+/+/ +

+/+

+/+

+/+

+/+

+/+

+/+

+/+

+/+

N/A

−

N/A

+/ +

+/ +

−/−

−/−

−/−

−/−

+

−

−

+

−

−

−/ + −

−/−

N/A

+

+

+

+

+

−

+

+

+

10.4

8.5

8.6

9.8

7.1

5.1

8.1

7.7

6.6

Total

(continued)

Masonry Force–displacement K Energy Damages wallets (fm,inf ; (envelopes, global (Kini , dissipation observed results) Em,inf ; Ksec ) fb,paralell,inf , fb,perpendicular,inf , Ss,inf )

2.7 Global Overview of OOP Tests on Infill Masonry Walls 15

Furtado et al [45] 3

24

+

+

+

+

+

+

Note + , Present; −, Absent; N/A, Not Applicable

4

Risi et al [44]

23

6

(Di Domenico et al. [42, 43]

+

+

+

+

+

+

+/±

+/±

+/±

+/+/±

−/−/ + / +

−/−/ + / +

Number Specimens’ Test Loading Frame Material properties of tests details setup Protocol detailing Mortar Masonry (type, units (% of fm,mo ; voids; fm,mo ) specific weight; fm,parallel,unit ; funit,b,perp )

22

Author and year

Table 2.1 (continued) Test results

+/+/+/+/ +

+ / ± /−/−

+ / ± /−/−

+/+

+/+

+/+

−/−

+/ +

+/ +

+

−

−

+

+

+

Masonry Force–displacement K Energy Damages wallets (fm,inf ; (envelopes, global (Kini , dissipation observed results) Em,inf ; Ksec ) fb,paralell,inf , fb,perpendicular,inf , Ss,inf )

9.8

8.5

8.5

Total

16 2 Out−of−Plane Behaviour of Masonry Infill Walls Experimental Tests: …

2.8 OOP Load Application Strategies and Protocol 100

138

17

67

69

50

46

33

23

17

0

0

115

83

92

67

69

50

46

33

23

17

0

0

Shake-Table

1

1:1.5

Type of Test

(a)

1:3

(b) 138

18.59%

Number of tests

As built with bed joints reinforcement As built Retrofitted

75.64%

1:2

Specimen Scale

138

#total:156

115

115

92

92

69

69

46

46

23

23

0

0

5.77%

(c)

Percentage of total tests (%)

Half-cyclic

Number of tests

92

Percentage of total tests (%)

Number of tests

83

Percentage of total tests (%)

#total:156

115

Monotonic

100

138

#total:156

RC

Steel

Type of Frame

(d)

Fig. 2.3 Global overview results: a Type of test; b Specimen’ scale; c Specimen condition; and d Type of frame

in the sub−Sect. 3.2. From the total amount of 153 tests that make up the database, it was found that 67% of them are composed by monotonic tests, 23% cyclic tests and, finally 9% shake−table tests (Fig. 2.3a). Regarding the specimens’ scale, 66% were full−scale tests while the remaining are related to scaled specimens, namely 1:1.5, 1:2, 1:3 and 2:3 (Fig. 2.3b). Most of the tests were carried out in as−built infill walls (76%), the remaining ones were on retrofitted panels (19%) and on panels with bed joints’ reinforcement (6%), as shown in Fig. 2.3c. It was observed that most of the panels tested were surrounded by an envelope frame made of RC instead of steel (Fig. 2.3d).

2.8 OOP Load Application Strategies and Protocol Regarding the quasi−static tests, different loading protocols were considered by the different authors. In fact, a 3D motion characterizes an earthquake, which means that the panel is subject to combined and multiple loadings, such as IP−OOP horizontal loadings combined with gravity loads (applied by the top structural frame members).

18

2 Out−of−Plane Behaviour of Masonry Infill Walls Experimental Tests: …

The laboratory reproduction of these combined multiple−directions is very difficult to simulate due to the test setup complexity. From the tests that available in this database, most of them (49%) are related to pure OOP tests. Concerning 15% of them, which are related to combined IP−OOP tests, the loading protocol was divided in two independent stages, namely: (i) first, an IP test was performed to impose some certain level of damage to the panel; (ii) second, a pure OOP loading was applied on the damaged infill panel. These tests are very important to understand the effect of the damage, due to previous IP loading demands, in the infill panel OOP capacity. The IP−OOP tests combined with the application of gravity load totalize 22% of the tests. Concerning this type of loading protocol, different strategies were assumed for the application of the gravity load application namely: (i) some authors applied it directly in the top of the upper beam (Figs. 2.4a and b); (ii) while some others applied it in the top of the adjacent columns (Figs. 2.4c and d). From the analysis of the literature studies, the most correct strategy to apply the gravity load is not consensual, due to uncertainties associated with the loading distribution, which can be transferred to the wall and and/or to the columns.

(a)

(c)

(b)

(d)

Fig. 2.4 Gravity load application approaches: a and b distributed load on the top beam (adapted from [36], c and d) Local loads on the top of the columns (adapted from [12])

2.8 OOP Load Application Strategies and Protocol

OOP: Pure OOP test OOP+IP: Combined OOP test with previous IP damage

100

124

90

112

80

99

80

87

70

74

60

62

50

50

40

37

30

25

20 10

OOP+G: Combined OOP test with gravity load

Number of tests

87

OOP+IP+G: Combined OOP test with previous IP damage and gravity load

70

74

60

62

50

50

40

37

30

25

20

12

10

12

0

0

0

OOP

OOP+IP

OOP+IP+G

Loading Protocol

(a)

OOP+G

100

#total:124

90

N.A. - Not aligned

Percentage of total tests (%)

99

#total:124

Number of tests

112

Percentage of total tests (%)

124

19

0 Airbag

4 points N.A.

4 points

8 points

OOP loading application

(b)

Fig. 2.5 Global overview results: a loading protocol; b OOP loading application strategy

Moreover, again the high level of complexity of the test setup required to apply the axial load and the strategy to apply the horizontal load, as already mentioned, is an important limitation. In a real scenario, the wall can be subjected to certain gravity load (infill panel vertical load), imposed by the short and long−term deformation of the top beam/slab, however it is very hard to quantify this accurately. Due to this, most of the experimental tests considered in this database were carried out without application of any additional vertical load. Finally, 14% of the tests were OOP tests combined with gravity load. Figure 2.5a show the number of tests performed according to each loading protocol. Over the literature, different strategies were adopted to apply of OOP load, from uniform distributed loadings throughout the entire panel to local loads specifically distributed. Starting from the first ones, this concept was first tested and validated by [46] that used water bags to apply an uniform distributed OOP loading in red bricks walls strengthened with fibre reinforced polymer (FRP) composites. In the same year, [47, 48] tested the same concept, but using airbags to apply uniform OOP loading to stone masonry walls. Griffith et al. [47, 48] carried out tests by using double−actuation airbag’ system, one in each face of the specimen which allowed to perform OOP cyclic, instead of the load−unload inherent to the adoption of only one actuation airbag system. One of the disadvantages related to the use of double−actuation airbag system is the impossibility of observation of the specimen damages’ evolution throughout the test. Apart of that, the synchronization between the two−airbag actuation components is also considered a difficulty. The use of local loads at four or eight points applied by hydraulic actuators is commonly adopted to overcome the difficulties inherent to the complexity of test setups that ensure the entire mobilization of the panel. However, some authors point out that one disadvantage concerning this approach is the possibility of introduction of a local failure or modification of the global infill panel behaviour when compared with the expected one when subjected to a real earthquake.

20

2 Out−of−Plane Behaviour of Masonry Infill Walls Experimental Tests: …

From the database tests, it is possible to observe that most of them were performed using airbags (about 69%) and the remaining ones by using 4 points not aligned approach (23%), 4 aligned points (6%) and 8 points (2%), as shown in Fig. 2.5b. Concerning the loading protocol, about 77% of the tests are monotonic tests and the remaining ones are semi−cyclic tests (load−unloading). From the analysis of all the testing campaigns, different loading protocols were adopted regarding the definition of number of half−cycles per peak displacements, the definition of the OOP peak displacements applied and finally, concerning the definition of the ultimate displacement. The use of airbags revealed to be very effective as proved by the recent works developed in the masonry field [47–49]. As can be observed in Fig. 2.5b, about ninety OOP tests were carried out within the application of the OOP loading with airbags. Silva [39] carried out OOP quasi−static tests using airbags that reacted against an independent steel reaction frame. The authors also restricted the OOP displacements of the envelope frame by using steel braces connected to the top beam. Alternatively, Lunn et al. [50] performed OOP quasi−static tests using airbags placed between the panel and an RC shear wall. This wall was attached to the envelope RC frame in 14 points (5 connections in the top beam, five connections in the bottom beam and two in each column) using steel rebars. A different approach was assumed by [40], where it was applied the OOP loading in 4 points not aligned by using a hydraulic actuator reacting against a steel structure and dividing the OOP load into four points. Finally, [10] performed full−scale OOP quasi−static tests by using a hydraulic actuator also reacting against a steel structure and distributing the OOP loading into four aligned points. Different masonry units were used to build the infill walls included in this database. The selection of each type is because they represent each country’s most common masonry unit. Clay units are the most common material selected to be the brick masonry units. A large part of the tests is related to infill panels made with Hollow Clay Horizontal Brick (herein designated HCHB) and Solid Clay Brick (SCB), respectively 31% and 29%. The remaining ones were made with Vertical Hollow Concrete Blocks (VHCB), Concrete Solid Brick (CSB) and Clay Vertical Hollow Brick (CVHB), respectively 16%, 10% and 14% (Fig. 2.6a). From the analysis of the specimens’ geometric dimensions, a large variability is visible, as shown in Fig. 2.6b. On the other hand, it can be observed that most of the infill panels tested have larger width than height and that only one specimen has a square dimension. Concerning the panel area, herein designated Ainf , it appears not directly related to the panel thickness as can be observed in Fig. 2.6c. From the plot, it was derived empirical Eq. 2.1, which is suggested to establish some relationship between the thickness and the specimens’ panel area. On the other hand, a trend for direct relation can be found between the panel aspect ratio (Hinf /Linf ) and the panel slenderness (Fig. 2.6d) and expressed by the fitted Eq. 2.2. However, due to the large dispersion of results, the obtained regression equations are characterized by their poor accuracy, as proved by the correlation coefficient R2 = 0.68 and R2 = 0.55, respectively. By excluding some outliers, the accuracy will certainly increase.

2.8 OOP Load Application Strategies and Protocol

21 5000

100

VHCB SCB HCHB CSB CVHB

80

VHCB - Vertical Hollow Concrete Block SCB - Solid Clay Brick HCHB - Horizontal Clay Hollow Brick CSB - Concrete Solid Brick CVHB - Clay Vertical Hollow Brick

4000

Panel Height Hinf (mm)

Percentage of total tests (%)

#total:156

60

40

40 tests

61 tests

20 22 tests 14 tests

3000

2000

1000

19 tests

0

0 VHCB

SCB

HCHB

CSB

0

CVHB

1000

2000

(a) 350

200

50

Slenderness Hinf / t

Panel thickness (mm)

250

4000

5000

(b) 60

VHCB SCB HCHB CSB CVHB

300

3000

Panel Lenght Linf (mm)

Masonry unit type

t=0.104xA0.4462 inf

150 100

40

VHCB SCB HCHB CSB CVHB Hinf /t=12.77xln (Hinf/Linf)+23.11

30

20

50

10 0 0

20

40

60

80

100

120

140

0.00

0.25

Panel Area Ainf x105(mm2)

0.50

0.75

1.00

1.25

1.50

1.75

2.00

Aspect Ratio Hinf / Linf

(R2 =0.68)

(R2 =0.55)

(c)

(d)

Fig. 2.6 Global overview results: a masonry unit type; b panel height versus panel length; c panel area versus panel thickness; d aspect ratio versus slenderness 0.4462 t = 0.104 × Ain f

) ( Hin f Hin f + 23.11 = 12.77 × ln t L in f

(2.1) (2.2)

Where Hinf , Linf , t and Ainf are the infill panel height, length, thickness and area, respectively.

22

2 Out−of−Plane Behaviour of Masonry Infill Walls Experimental Tests: …

2.9 OOP Behaviour of As-Built Infill Masonry Walls 2.9.1 Effect of RC Frame Type and Detailing Different RC frame reinforcement detailing approaches were found over the different experimental studies. Some authors assumed the design practice according to past codes, where the seismic action is neglected and the reinforcement detailing is very poor, with reduced amount of transverse reinforcement, poor beam−column joint detailing and with insufficient anchorage length [12, 39]. However, these represent an important part of the building stock of each country. For example [12], developed a statistical study concerning the Portuguese RC buildings and found that more than 50% of the buildings in 2015 were designed according to codes that consider very low and insufficient seismic provisions. By contrast, some others assumed the design and reinforcement detailing according to the most recent national and international codes, where the expected seismic action and design of high ductility frame members are considered. Concerning this last approach, special attention was given the authors to the beam−column joints, especially to the top and bottom part of the columns [11, 32]. As shown in Fig. 2.7a, 40% of the RC frames used to test the OOP capacity of the infill panels were designed with high ductility demand and 26% with low ductility. The authors did not explain the reinforcement design criteria of the remaining 33% of the specimens. The maximum OOP strength capacity of 51 specimens, for which the reinforcement design was divided according to high or low ductility, is plotted in Fig. 2.7b. Due to a large number of variables among the tests, it is impossible to assess the possible effect of the frame detailing in the panel OOP capacity. From the point of view of pure OOP tests, the frame detailing is not expected to play an important role since the frame is not being tested. Only the infill panels 70

54

40

46

50

38

40

31

30

23

20

15

10

8

0

0 High Ductility

Low Ductility

RC frame detailling

(a)

N/A

Percentage of total tests (%)

60

Maximum OOP Strength FOOP,max (kN/m2)

Number of tests

#total: 146

35

#total: 51 Tests with complete reinforcement detailling available and without previous IP drift

30 25 20 15 10 5 0

Low ductility High ductility RC frame detailling

(b)

Fig. 2.7 As built specimens’ results: a RC frame detailing; b Maximum OOP strength capacity according to RC frame detailing

2.9 OOP Behaviour of As-Built Infill Masonry Walls

23

are subjected to the OOP loadings. According to the literature [4], the frame can influence and affect the OOP performance of the panel only in terms of adhesion conditions provided by the interface joints between frame concrete and masonry above the panel contour. On the other hand, different considerations can be drawn in the case of combined IP−OOP tests where the frame detailing can play an important role, since the panel damage caused by the IP loadings depends directly of the infilled frame global behaviour. For example, an infilled RC frame designed according to code standards with high seismic demands will perform better when subjected to lateral loadings, which probably leads to lower levels of damage in the panels. On the other hand, an RC frame not designed according to the most up to date codes, without proper detailing, it is expected that high levels of damage occur, which can affect significantly the panels and reduce their OOP capacity. Thus, it should be pointed out that, in a real earthquake scenario, the RC frame reinforcement detailing can contribute to the infill OOP behaviour.

2.9.2 Effect of Geometric Dimensions and Mechanical Properties The effect of the infill panel geometric dimensions (such as aspect ratio and slenderness) and their mechanical properties on the maximum OOP strength will be assessed in this subsection. The results were grouped according to each masonry unit, namely: hollow clay horizontal bricks (HCHB), solid clay brick (SCB), vertical hollow concrete block (VHCB), concrete solid block (CSB) and clay vertical hollow brick (CVHB). For this analysis, only tests without gravity load and openings were considered. From the relationship between the panel aspect ratio and the maximum OOP strength, plotted in Fig. 2.8a, it is possible to observe a reduction of the panel maximum OOP strength with increased aspect ratio. The empirical relationship (given by Eq. 2.3) is proposed herein to estimate the correlation between the panel aspect ratio and the corresponding maximum strength. The accuracy of the proposed equation is affected by the large variability of results, as can be observed by the correlation coefficient equal to R2 = 0.62. Relatively to the effect of the panel slenderness (global trends), as expected, it seems that the panel maximum OOP capacity slightly reduces with the increase of the slenderness (Fig. 2.8b). However, this analysis is affected by the dispersion of results. For a better understanding of the effect of this parameter, the analysis was performed group by group. For instance, two empirical equations are provided for the trends observed for the groups VHCB and HCHB (Eqs. 2.4 and 2.5, respectively) which led to correlation coefficients of R2 = 0.74 and R2 = 0.87. From this analysis, only the panels with four edges constrained were considered. These results indicate

2 Out−of−Plane Behaviour of Masonry Infill Walls Experimental Tests: …

VHCB SCB CSB HCHB CVHB

50

40

Eq. 3 13

Maximum OOP strength FOOP,max (kN/m2)

2 Maximum OOP strength FOOP,max (kN/mm (kN/m2))

24

#total:41

30

20

10

0 0.0

0.5

1.0

1.5

2.0

2.5

VHCB SCB CSB HCHB CVHB

60

50

#total:41

40

Eq. 15 5 30

4 Eq. 14

20

10

0

3.0

0

Aspect Ratio Hinf/W inf (R2=0.62)

10

(R2=0.74 –

20

VHCB HCHB

60

#total:16 50

40

30

Excluded

6 Eq. 16

10

0 0.4

0.5

0.6

0.7

0.8

0.9

50

#total:43 40

30

20

10

0 0.0

0.5

Maximum OOP strength FOOP,max (kN/m2)

Maximum OOP strength Fmax (kN/m2)

40

30

Eq. 18 8 Excluded

10

0.1

0.2

0.3

0.4

0.5

0.6

Masonry wallets flexural strenght paralell to bed joints fb,par (MPa)

(R2=0.70) (e)

1.5

2.0

2.5

3.0

3.5

(R2=0.49) (d)

#total:35

0 0.0

1.0

Masonry wallets compressive strenght fm (MPa)

VHCB SCB HCHB CVHB

20

Excluded

7 Eq. 17

(R2=0.57) (c)

50

– VHCB)

VHCB SCB HCHB CVHB

60

Percentage of voids (%)

60

40

(b) Maximum OOP strength FOOP,max (kN/m2)

Maximum OOP strength FOOP,max (kN/m2)

(a)

20

30

Slenderness Hinf/t HCHB and R2=0.87

60

50

40

SCB HCHB CVHB

#total:22 Excluded

30

20

Eq. 19 9

10

0 0.00

0.25

0.50

0.75

1.00

1.25

1.50

1.75

2.00

Masonry wallets flexural strenght perpendicular to bed joints fb,perp (MPa)

(R2=0.68) (f)

Fig. 2.8 As built specimens’ maximum OOP strength: Assessment of the effect a aspect ratio; b slenderness; c percentage of voids; d masonry wallets compressive strength; e flexural strength parallel to the horizontal bed joints; and f perpendicular to the horizontal bed joints

that the maximum strength of the VHCB group decrease slightly for larger panel’ slenderness when compared with the HCHB group. The particular analysis of the slenderness effect will be discussed later in this section. Concerning the results of groups CVHB and SCB, no conclusions can be achieved due to the large dispersion of results. It is necessary to increase the number of experimental tests for each group of masonry units to better assess the effect of

2.9 OOP Behaviour of As-Built Infill Masonry Walls

25

this parameter, since many authors point out that the slenderness is one of the most important variables that affect the infill wall OOP behaviour. The reduced amount of data containing information regarding the percentage of voids limited the analysis of this effect in the panel OOP capacity. However, the reduced data (16 elements) still suggests that the increase of the percentage of voids leads to lower maximum OOP strength (according to Eq. 2.6), which relates to the reduced capacity of the panel to develop arching mechanism and to the crushing of bricks associated with this mechanism (Fig. 2.8c). ( Fmax = 17.985 ×

Hin f L in f

Fmax,V H C B = 17.559 × ex p

(2.3)

(H ) −0.023× int f

Fmax,H C H B = 17.792 × ex p Fmax = 3.4369 × (%

)−1.377

(H ) −0.057× int f

voids )

−2.031

(2.4) (2.5) (2.6)

Where Fmax , Fmax,VHCB , Fmax,HCHB are the maximum OOP strength of as−built panels, as−built panels made with VHCB and HCHB units, respectively. Hinf and t are the infill panel height and thickness, respectively. Finally, %voids is the masonry unit percentage of voids. According to the tests available in this database, the masonry properties play an important role in the panel OOP capacity. For example, the maximum OOP strength increase with the masonry compressive strength (Fig. 2.8d). From the plot, an increase of the masonry compressive strength from 0.5 to 3 MPa yields an increment of 200% of the panel maximum OOP strength, however some variation of the results were observed. The Eq. 2.7 was derived with a correlation coefficient (R2 ) of 0.49. To avoid lower correlation, some results were excluded concerning this effect. Regarding the effect of the masonry flexural strength parallel to the horizontal bed joints, shown in Fig. 2.8e, the OOP maximum strength increases for larger flexural strength. Even though the flexural strength varied slightly (between 0.1–0.55 MPa), this variation increased the OOP strength capacity of the panel around 200%. Equation 2.8 is proposed to predict this parameter’s effect and suggested a good correlation coefficient equal to R2 = 0.70. In the same way, the flexural strength perpendicular to the horizontal bed joints appears to have the same effect, since it can be observed that the maximum OOP strength increases for larger flexural strength (Fig. 2.8f). From these results, Eq. 2.9 is proposed to estimate the relationship between the panel maximum OOP strength and their flexural strength perpendicular to the horizontal bed joints with a correlation coefficient R2 = 0.68. From this analysis, it is impossible to achieve any conclusion regarding the effect of the panel aspect ratio, compressive strength and flexural strength (parallel to the

26

2 Out−of−Plane Behaviour of Masonry Infill Walls Experimental Tests: …

horizontal bed joints) in the panels made with VHCB units. Further tests are needed to achieve solid conclusions. Finally, from the database analysis, only a few authors provided information about the masonry units’ compressive strength. Thus, it was not possible to assess the effect of this parameter. Fmax = 0.5467 + 16.122 f m − 1.4572 f m2 + 0.5467 f m3

(2.7)

Fmax = −26.616 + 136.02 × f b, parallel

(2.8)

Fmax = −2.1168 + 28.277 × f b, per pendicular

(2.9)

Some authors studied the effect of the slenderness ratio on the OOP capacity of the infill panels, for example Di Domenico et al. [42, 43] carried out a testing campaign comprised by specimens made with HCHB units with two different thicknesses (80 and 120 mm) and three different boundary conditions (4 edges constrained, 3 edges and 2 edges constrained). From the force–displacement results the authors found that panels with thicker masonry units (high slenderness ratio) evidenced a reduction of the OOP strength capacity around 40–50%. Recently, [41] tested two scaled specimens built with HCHB units with two different thicknesses (60 and 80 mm). The main aim was to test a “typical” internal and an external leaf of a building façade, SIF−O−1L−A and SIF−O−1L−B. The force–displacement results showed a reduction of about 12% of the maximum strength, lower deformation capacity and a reduction of energy dissipation capacity by about 20% for the same OOP displacement level and about 50% regarding the total cumulative energy dissipation. The cracking pattern was modified since the thicker specimen behaved with a trilinear cracking and the larger one as a five hinge crack. The relationship between the OOP drift at the maximum strength (dFmax ) and the drift at 80% of the maximum strength (d0.80Fmax ), at the post−peak stage, and the drift at the ultimate strength (dFult ), is plotted in Figs. 2.9a and b, respectively. From that, it is observed that the drift at 80% of the maximum strength is around 1.2 to 3 times the value of dFmax . It can be observed that panels made with HCHB units have a larger deformation capacity to reach larger OOP drifts after having reached the maximum strength. The same is observed when dFmax is compared with the dFult . A variation of dFult between 1 and 3.5 times the corresponding drift for maximum peak load (dFmax ) for panels with HCHB units, 1.5 to 2.5 times for panels made with CSB units and 1.1 to 4 times for panels made with VHCB units. Concerning the secant stiffness variation, the relative stiffness was calculated based on the secant stiffness at three different levels in the pre−peak phase, namely: (i) 30% of the maximum peak load (Ksec,0.3Fmax ); (ii) 60% of the maximum peak load (Ksec,0.6Fmax ); and (iii) at the maximum peak load (Ksec ).

2.9 OOP Behaviour of As-Built Infill Masonry Walls

27

(a)

(b)

(c)

(d)

Fig. 2.9 As built specimens’ results: a OOP driftFmax versus OOP drift0.8Fmax ; b OOP driftFmax versus OOP driftFult , and c Relative stiffness; d OOP driftFmax versus panel aspect ratio; e OOP driftFmax versus panel slenderness

From the results plotted in Fig. 2.9c, a large decrease of the relative stiffness for OOP drift levels lower than 0.5% can be observed. For example, for almost 40% of the data, the specimen stiffness decreased more than 50%. For OOP drift values higher than 0.5%, a smoother reduction of the relative stiffness is visible. The relationship between the panel aspect ratio and the OOP drift at the maximum peak load was assessed (Fig. 2.9d) and the results show large dispersion. However, the results indicate a slight increase of dFmax for larger panel aspect ratio. In the same way, the relationship between the panel slenderness and the OOP drift at the maximum peak load was assessed and plotted in Fig. 2.9e. The result shows the increase of dFmax for lower panel slenderness.

2.9.3 Effect of Gravity Load From the literature, only two different experimental studies carried out OOP tests in as−built specimens considering the vertical dead load.

28

2 Out−of−Plane Behaviour of Masonry Infill Walls Experimental Tests: …

The first test date from 1994, where [4] tested one specimen specifically to assess the influence of the gravity load (in the RC adjacent columns) in the panel response. A normalized axial stress of 0.04 was applied at each column’s top. The force– deflection curve between the specimens with and without gravity loads was very similar. The authors reported that new cracks were not visible during the gravity load test. In 2011 [31], carried out a set of OOP tests in IM walls combined with gravity load distributed along the top frame beam. However, the major goal of the testing campaign was not to assess the gravity load effect; the authors did not test any specimen without gravity load.

2.9.4 Effect of Openings and Panel Support Conditions The effect of openings in the panel OOP response was only studied in 6 tests. The first one was performed by Dawe et al. [2, 3], in which it was tested an infill masonry wall with a central opening from which they concluded that no significant decrease of the panel maximum strength was observed, but some differences occurred during the post−peak phase. The authors concluded that the central opening affected the ability of the central strips to support the OOP load and reduced the post−peak phase deformation capacity. The opening reduced also the OOP drift corresponding to the maximum peak load. The authors also concluded that an opening with small dimensions does not affect significantly the strength provided by the arching mechanism. Later, in 2014 [10] tested two specimens partially filled, namely: (i) specimen TA4 with a central opening, and (ii) specimen TA5 with a vertical infill stripe configuration. The test TA4 was subjected first to an IP cyclic test up to a maximum drift of 1% and after that it was subjected to OOP half−cyclic loads applied in four aligned points. On the other hand, the specimen TA5 was only subjected to half−cyclic OOP loadings aiming at reproducing single−bending along the vertical direction. The authors did not test any reference specimen without openings. In addition, due to the large influence of the previous damage and the unusual configuration of the wall TA5, no conclusions were drawn by the authors regarding the openings effect. Preti [51] tested one specimen with a central opening where engineered sliding joints were used. The loading protocol was divided into two stages, first the application of a cyclic IP loading and then the application of half−cyclic OOP loadings. Due to the use of the sliding joints, this test is excluded from this analysis of the effect of the opening. To finish, Akhoundi et al. [37, 38] carried out an OOP test of a scaled infill panel with a central opening. The authors observed that the panel collapsed for low OOP displacements. The maximum deformation achieved was about 25% of that obtained by the reference specimen. A reduction of the panel initial stiffness was also observed around 13% (similar to the percentage of the opening area).

2.10 Assessment of the Previous IP Damage Effect on the Infill Walls OOP …

29

Concerning the assessment of the panel’ width support effect in the panel OOP capacity it was not found in any work in which this variable was studied. The only test available in the literature is one carried out in this thesis that will be described later.

2.10 Assessment of the Previous IP Damage Effect on the Infill Walls OOP Capacity 2.10.1 Background The damages caused by prior IP demands affect the infill walls OOP performance, namely causing the reduction of the maximum strength capacity and modifying the expected failure mechanism. From the database tests, eight works studied this IP−OOP interaction by performing OOP tests on panels that were first subjected to IP tests and different IP damage levels. The first author that studied this effect was [4], who performed three tests (specimens 1, 2b and 3) in infill panels made with SCB units (slenderness = 33.9). The reference specimen 1 was subjected only to a uniform monotonic OOP loading and reached a maximum peak load of 8.18kN/m2 . On the other hand, specimens 3 and 2b were tested first to IP loading demands to different drift levels (0.22% and 0.34% respectively) and, afterwards, subjected to monotonic OOP loading until the panel collapse. The maximum strength achieved by each one was 5.98kN/m2 and 4.02 kN/m2 , respectively, corresponding to a reduction of 27% and 51% relative to the reference one. The authors provided no details regarding the failure mechanism. In 2001, Calvi et al. [7] tested two full−scale single panels made with HCHB units, with slenderness equal to 20.4, subjected first to an IP cyclic loading and later to a OOP monotonic loading. The specimens 6 and 2 were first subjected to a maximum IP drift equal to 0.4% and 1.2%, respectively. Afterwards, both were subjected to a monotonic OOP load applied in four not aligned points. The tests were compared with the reference specimen 10, which was only subjected to a monotonic OOP loading. The results showed a reduction of the OOP maximum strength about 75% and 82%, respectively. However, no reports were provided concerning the cracking pattern or failure mechanism of any of them. Pereira [8] carried out a series of 7 tests, 1:1.5 scaled, made with HCHB units, with slenderness equal to 11.3 that were subjected to combined IP−OOP tests. Nevertheless, no specimen was subjected only to pure OOP loading, thus, the effect of the previous IP damage in the panel response was not assessed. The major goal of the testing campaign was to discuss the efficiency of different retrofit solutions. A similar testing campaign was carried out by [32] who tested 6 full−scale panels made with HCHB units, with slenderness equal to 13.8, with different types of mortar and strengthening solutions. The authors again did not discuss the effect of the previous IP damage in the panel OOP capacity.

30

2 Out−of−Plane Behaviour of Masonry Infill Walls Experimental Tests: …

In 2015, Francesca [11] tested 8 full−scale specimens with the same geometric characteristics as those tested by [32] to assess the efficiency of strengthening strategies. The authors tested panels made with HCHB units characterized by high percentage of voids. Besides the fact that it was not tested with any reference specimen, one of the conclusions drawn by the authors was that the OOP response of panels with previous damage due to IP demands is very fragile and with unstable post−peak behaviour. In 2018, [40] conducted a testing campaign composed of 3 combined IP−OOP tests on 2:3 scaled IM walls made with HCHB units, with slenderness equal to 20.3. The specimens IP_OOP_L, IP_OOP_M and IP_OOP_H were subjected to a maximum IP drift equal to 0.16%, 0.37% and 0.58%, respectively. After that, all of them were subjected to a monotonic OOP loading. These specimens were compared with the reference specimen OOP_4E which was not subjected to any previous IP loading. The authors found a reduction of the specimens’ maximum strength (IP_OOP_M and IP_OOP_H) by about 53% and 74%, respectively. An unexpected result was achieved by panel IP_OOP_L which evidenced an increase of the maximum strength by around 6%. The authors justified this result with experimental variability. Ricci et al. [52, 53] tested four scaled infill panels, built with HCHB units 120 mm thick (slenderness ratio equal to 15.2) and the main aim was to assess the effect of different levels of prior damage and to compare with the results obtained by [40] in which specimens were used the same aspect ratio but with thicker panels (HCHB units 80 mm thick). The specimens 120_IP + OOP_L, 120_IP + OOP_M and 120_IP + OOP_H were subjected to a maximum IP drift equal to 0.21%, 0.50% and 0.89%, respectively. These specimens were compared with the reference one 120_OOP_4E. From the results, the authors pointed out that all the specimens showed an almost equal initial stiffness and, an almost similar strength (when reached) roughly equal to 120kN. Concerning the maximum peak load, the specimen 120_IP_OOP_L reach a similar value. On the other hand, the remaining specimens 120_IP_OOP_M and 120_IP_OOP_H decreased about 33 and 45% their maximum strength capacity. Akhoundi [41] performed four experimental tests on scaled specimens, built with HCHB units, three of which were subjected to prior IP test with different drift levels. The result of the reference specimen SIFO−1L−B was than compared with the specimens SIF−IO(0.3%)−2L(NC)−B, SIF−IO(0.5%)−2L(NC)−B and SIF−IO(1%)−2L(NC)−B that were subjected to a prior IP test with a maximum drift equal to 0.3%, 0.5% and 1%, respectively. The results revealed that the OOP strength capacity reduced with the increase of the prior IP drift demand, namely a reduction of about 15, 34 and 49%. It was also observed that the energy dissipation capacity reduced about 13, 40 and 65% for the same OOP displacement demand. To conclude [44], carried out a testing campaign comprised by four specimens, one reference specimen designed OOP (pure OOP test), and three specimens that were first subjected to prior IP test with three different IP drift demands, namely 0.15% (IPL−OOP), 0.28% (IPM−OOP) and 0.51% (IPH−OOP). From the tests, the authors observed a reduction of the OOP strength due to IP damage equal to 24% for the panel IPM−OOP, equal to 35% for the panel IPH−OOP and about 2% for

2.10 Assessment of the Previous IP Damage Effect on the Infill Walls OOP …

31

the panel IPL−OOP. The authors pointed out that similar trends were visible in other parameters, such as secant stiffness at cracking stage, secant stiffness at peak load). In the following sub−sections, a detailed analysis is presented concerning the effect of prior damage due to different levels of IP drift demands in the OOP capacity of the panels along the different stages such as first−macro cracking, and maximum peak load and ultimate stage. All the results of the damaged specimens are compared with the panels tested under pure OOP loadings.

2.10.2 Cracking Strength and Secant Cracking Stiffness Some authors provided information regarding the instant formation of the first major visible cracking. The corresponding strength (here designated cracking strength, FOOP, crack ), OOP drift and (dOOP, crack ) and secant cracking stiffness (KOOP, crack,sec ) were collected. The results obtained by [4, 12, 40], Ricci et al. [52, 53], [41] and [44] were collected. The ratios between the reference specimens and the panels with previous IP damage were computed to assess the impact of the previous damage due to different IP drift levels in the panel OOP response. The results are summarized in Table 2.2. As expected, it is observed that all the specimens reached lower cracking strength for higher levels of previous IP drift (Fig. 2.10a). The empirical Eq. 2.10, based on the experimental results, that correlate the ratio between the damaged and undamaged strength is proposed with good accuracy (R2 = 0.68). The accuracy (herein designated ACC) of the equation proposed was assessed for all the tests, as shown in Fig. 2.10b. This assessment found an average ACC equal to 1.19, a coefficient of variation (C.o.V.) equal to 44% and a standard deviation equal to 0.53. To achieve a better ACC the specimen Inf_11 and IP_OOP_H results were considered as outliers. It is recalled that ACC represents the ratio between the prediction, according to the empirical equation proposed, and the experimental result obtained by each specimen, which allows to assess the equation’s accuracy. Regarding the secant cracking stiffness, the same trend is observed (as shown in Fig. 2.10c), however one difference was found, namely the specimen IP_OOP_L achieved a secant stiffness 20% larger than the reference specimen, which is a result that is clearly out of the trend presented by the other results presented. Equation 2.11 aims to correlate the ratio between damaged and undamaged secant cracking stiffness due to different previous IP drift levels (Fig. 2.10d). Lower ACC was achieved and it can be observed that this proposal overestimates the cracking secant stiffness ratio. The ACC of this equation resulted in a mean value equal to 1.09, which is a good approximation, and C.o.V. equal to 46%. (

FO O P,crack,damaged FO O P,crack,undamaged

)

= 0.3554 × I P −0.395 pr ev

(2.10)

Ricci et al. [52, 53]

Di Domenico [40]

Furtado [12]

0.21 0.50 0.89

120_IP + OOP_L

120_IP + OOP_M

120_IP + OOP_H

0.60

0.58 0 (Ref)

IP_OOP_H

120_OOP_4E

1.98

0.37

IP_OOP_M

4.51

2.60

6.12

9.67

6.39

2.93

0 (Ref) 0.16

2.39

OOP_4E

0.3

Inf_11

2.34

1.86

5.17

IP_OOP_L

0.5 0 (Ref)

Inf_03

Inf_08

0 (Ref)

Inf_02

1.89 3.20

0.22 0.34

3

2b

0.74

0.41

0.96

1.51

N/A

0.13

0.44

0.65

N/A

1.02

N/A

0.36

N/A

0.44

1.61

0.79

1.25

4.29

11

0.28

0.73

7.00

5.88

2.99

17.87

7.83

20.75

1.04

0.07

0.11

0.39

N/A

0.05

0.12

1.19

N/A

0.17

N/A

0.28

N/A

0.20

0.31

N/A

(continued)

Dam/refer

Test result

N/A

5.18

Test result

Dam/refer

Secant Cracking Stiffness KOOP,crack,sec (kN/mm)

Cracking Strength FOOP,crack (kN/m2 ) 4.31

1

Angel [4]

Previous IP drift (%)

0 (Ref)

Specimen

Author

Table 2.2 Effect of previous IP damage: cracking strength and secant cracking stiffness

32 2 Out−of−Plane Behaviour of Masonry Infill Walls Experimental Tests: …

N/R

3.47

0.51

IPH −OOP

4.54

0.28

IPM −OOP

4.82

0 (Ref) 0.15

OOP

6.01

N/R

N/R

N/A

0.58

0.76

0.80

N/A

N/A

N/A

1.46

2.95

4.61

9.62

N/R

N/R

N/R

N/R

0.15

0.31

0.48

N/A

N/A

N/A

N/A

N/A

Dam/refer

Test result

N/A

Dam/refer

Test result N/R

Secant Cracking Stiffness KOOP,crack,sec (kN/mm)

Cracking Strength FOOP,crack (kN/m2 )

IPL −OOP

1

SIF−IO(1%)−2L(NC)−B

N/A: Not applicable N/R: Result not provided by the author

Risi et al [44]

0.3 0.5

SIF−IO(0.3%)−2L(NC)−B

0 (Ref)

SIFO−1L−B

Akhoundi [41]

SIF−IO(0.5%)−2L(NC)−B

Previous IP drift (%)

Specimen

Author

Table 2.2 (continued)

2.10 Assessment of the Previous IP Damage Effect on the Infill Walls OOP … 33

34

2 Out−of−Plane Behaviour of Masonry Infill Walls Experimental Tests: … 1.50

Angel et al. (1994) Furtado et al. (2016) Domenico et al. (2018) De Risi et al. (2019)

1.0

Ratio FOOP,crack,damaged/FOOP,crack,undamaged

Ratio FOOP,crack,damaged/FOOP,crack,undamaged

1.2

0.8

0.6

Eq. 20 10

0.4

0.2

0.2

0.4

0.6

0.8

Accuracy: mean:1.19 SD: 0.53 C.o.V.:44%

1.25 ACC=1.78

1.00

ACC=2.05 ACC=1.10

0.75

ACC=1.06 ACC=1.29

ACC=1.14 ACC=0.89

ACC=1.25

ACC=0.29

0.50

ACC=0.77 ACC=0.84

ACC=0.81

0.25

0.00

0.0 0.0

ACC=2.29

Test Result Calculated

1.0

2b 3 Inf_03 Inf_11 OOP OOP OOP 120_OOP 120 IP 120 IP 120 IP IPL IPM IPH L M H OOP_L OOP_M OOP_H OOP OOP OOP 4E

Specimens

Previous IP drift (%)

(R2 =0.68) (a)

(b) 1.50

Angel et al. (1994) Furtado et al. (2016) Domenico et al. (2018) De Risi et al. (2019)

1.0

Ratio KOOP,crack,damaged/KOOP,crack,undamaged

Ratio KOOP,crack,damaged/KOOP,crack,undamaged

1.2

0.8

0.6

Eq. 20 11

0.4

0.2

0.0 0.0

1.25

0.4

0.6

0.8

1.0

mean:1.09 SD: 0.50 C.o.V.:46%

ACC=2.04

1.00

0.75 ACC=0.75

0.50 ACC=0.97

ACC=0.82

0.25

ACC=1.15

ACC=2.29 ACC=0.69 ACC=0.96

ACC=0.65 ACC=0.50

0.00

0.2

Accuracy:

Test Result Calculated

ACC=0.90

ACC=1.26

ACC=1.26

2b 3 Inf_03 Inf_11 OOP OOP OOP 120_OOP 120 IP 120 IP 120 IP IPL IPM IPH H OOP_L OOP_M OOP_H OOP OOP OOP L M 4E

Specimens

Previous IP drift (%)

(R2 =0.74) (c)

(d)

Fig. 2.10 Effect of previous IP damage: a Cracking strength; b Assessment of the Eq. 2.10 accuracy; c Secant cracking strength; and d Assessment of the Eq. 2.11 accuracy

(

K O O P,crack,damaged K O O P,crack,undamaged

)

= 0.0473 × I P −1.37 pr ev

(2.11)

2.10.3 Maximum Strength and Secant Stiffness The specimens’ maximum strength (FOOP,max ) from the tests performed by [4], Calvi et al. [7, 12] and Ricci et al. [52, 53] were collected in order to compute some ratios between the reference specimens and the panels that were first subjected to a previous IP test. The same procedure was adopted for the secant stiffness (KOOP,sec ). The maximum peak strength and secant stiffness ratios between the results obtained by the undamaged and damaged panels are plotted in Figs. 2.11a and b, respectively, in function of the corresponding previous IP drift that they were subjected to.

2.10 Assessment of the Previous IP Damage Effect on the Infill Walls OOP … 1.2

Angel et al. (1994) Furtado et al. (2016) Domenico et al. (2018) Ricci et al. (2018) Akhoundi et al. (2018) De Risi et al. (2019)

1.0

Ratio KOOP,max,damaged/KOOP,max,undamaged

Ratio FOOP,max,damaged/FOOP,max,undamaged

1.2

0.8

Eq. 22 12 0.6

0.4

0.2

0.0 0.00

0.25

0.50

1.00

Angel et al. (1994) Furtado et al. (2016) Domenico et al. (2018) Ricci et al. (2018) Akhoundi et al. (2018) De Risi et al. (2019)

1.0

0.8

0.6

13 Eq. 23

0.4

0.2

0.0 0.00

1.25

0.25

0.50

0.75

1.00

Previous IP drift (%)

Previous IP drift (%)

(R2=0.78)

(R2=0.81)

(a)

(b)

1.25

Ratio FOOP,max,damaged/FOOP,max,undamaged

0.75

35

Test result Calculated

1.25

Accuracy: mean:0.97 SD: 0.33 C.o.V.:34%

ACC=0.98

ACC=0.95 ACC=0.90

1.00 ACC=1.19

ACC=0.86 ACC=0.99

0.75

ACC=1.05

ACC=1.34

ACC=0.74

ACC=1.31

ACC=0.75

ACC=0.75

ACC=0.42 ACC=0.35

ACC=1.37

0.50

ACC=1.58

0.25

0.00

2

3b

Inf_11

Inf_03

IP+OOP IP+OOP M L

IP+OOP H

120_IP OOP_L

120_IP OOP_M

120_IP SIF-0.3% SIF-0.5% SIF-1% IPL-OOP OOP_H

IPM-OOP IPH-OOP

Specimens

(c)

Ratio Ksec,max,damaged/Ksec,max,undamaged

1.25

Test result Calculated

Accuracy: mean:1.09 SD: 0.56 C.o.V.:51%

1.00 ACC=0.69

ACC=0.51

0.75 ACC=0.93

0.50 ACC=1.30

ACC=0.81

ACC=0.90 ACC=0.69

ACC=0.95

ACC=1.01

ACC=2.53

0.25

ACC=0.61 ACC=0.97

ACC=2.41

ACC=1.21

ACC=0.86

ACC=1.15

0.00

2

3b

Inf_03

Inf_11

IP+OOP IP+OOP M L

IP+OOP H

120_IP OOP_L

120_IP OOP_M

120_IP SIF-0.3% SIF-0.5% SIF-1% IPL-OOP OOP_H

IPM-OOP IPH-OOP

Specimens

(d)

Fig. 2.11 Effect of previous IP damage: a Maximum strength, b Secant Strength; c Assessment of the Eq. 2.12 accuracy, and d Assessment of the Eq. 2.13 accuracy

The results show a clear reduction of the maximum OOP peak load for larger levels of previous IP drift demand. A logarithmic relationship was found to follow all the test results with a good correlation (R2 = 0.78), as indicated in Fig. 2.11a, and it is given in Eq. 2.12. Comparing with the previous parameters, the results for the maximum strength ratio prediction were obtained with good accuracy, as proved by the value of ACC equal to 0.97 and a COV equal to 33%. Specimens 6, 2 and Inf_03 contributed significantly to the C.o.V. increment (Fig. 2.11c) (Table 2.3).

Ricci et al. [52, 53]

Di Domenico [40]

0 (Ref) 0.5 0 (Ref) 0.3 0 (Ref) 0.16 0.37 0.58 0 (Ref) 0.21 0.50 0.89

Inf_03

Inf_08

Inf_11

OOP_4E

IP_OOP_L

IP_OOP_M

IP_OOP_H

120_OOP_4E

120_IP + OOP_L

120_IP + OOP_M

120_IP + OOP_H

0.34

2b

Inf_02

0.22

3

Furtado [12]

0 (Ref)

1

Angel [4]

Previous IP drift (%)

Specimen

Author

Table 2.3 Effect of previous IP damage: Maximum peak strength and secant stiffness

5.37

6.49

9.67

9.74

1.37

2.44

5.44

5.12

3.18

4.57

1.86

5.17

4.02

5.98

0.55

0.67

0.99

N/A

0.27

0.48

1.06

N/A

0.69

N/A

0.36

N/A

0.49

0.73

1.04

1.06

4.29

5.24

0.23

0.33

3.44

4.07

0.59

1.74

7.83

20.75

0.89

1.03

3.33

0.19

0.20

0.82

N/A

0.07

0.10

0.85

N/A

0.34

N/A

0.28

N/A

0.27

0.31

N/A

(continued)

Dam/refer

Test result

N/A

Dam/refer

Test result 8.18

Secant Stiffness KOOP,sec (kN/mm)

Maximum Strength FOOP,max (kN/m2 )

36 2 Out−of−Plane Behaviour of Masonry Infill Walls Experimental Tests: …

5.73

0.51

N/A: Not applicable N/R: Result not provided by the author

IPH −OOP

6.71

0.28

9.39

IPM −OOP

8.79

5.13

0.15

1

SIF−IO(1%)−2L(NC)−B

6.68

8.60

0 (Ref)

0.5

SIF−IO(0.5%)−2L(NC)−B

IPL −OOP

0.3

SIF−IO(0.3%)−2L(NC)−B

0.65

0.76

1.07

N/A

0.51

0.66

0.85

1.07

1.95

3.02

4.93

0.33

0.56

1.17

3.61

0.22

0.40

0.61

N/A

0.09

0.16

0.32

N/A

Dam/refer

Test result

N/A

Dam/refer

Test result 10.07

Secant Stiffness KOOP,sec (kN/mm)

Maximum Strength FOOP,max (kN/m2 )

OOP

0 (Ref)

SIFO−1L−B

[41]

Akhoundi [44]

Previous IP drift (%)

Specimen

Author

Table 2.3 (continued)

2.10 Assessment of the Previous IP Damage Effect on the Infill Walls OOP … 37

38

2 Out−of−Plane Behaviour of Masonry Infill Walls Experimental Tests: …

The ratios between the damaged and undamaged specimens concerning the panels’ secant stiffness led to similar results (Fig. 2.11c). However, the ACC of the proposed Eq. 2.13 is limited by specimen IP_OOP_L result that reduced the correlation coefficient to R2 = 0.81. Finally, the mean ACC of Eq. 2.13 is equal to 1.09, but C.o.V. and SD increased to 51% and 0.56, respectively. The assessment of the Eq. 2.13 accuracy is shown in Fig. 2.11d. (

) FO O P,max,damaged = −0.458 × ln(I Ppr ev ) + 0.1764 FO O P,max,undamaged ( ) K O O P,max,damaged −0.897 = 0.1037 × I Ppr ev K O O P,max,undamaged

(2.12) (2.13)

2.10.4 IP−OOP Ultimate Displacement There are no recommendations over the international codes about the IP−OOP collapse interaction. Only some considerations can be found specifically for IP or OOP limit drifts independently but not with interaction. FEMA-274 [54] indicates that the Immediate Occupancy Performance Level is not necessarily related to initial cracking of the panel. Some cracking can be tolerated for typical occupancy conditions. Life Safety is related to extensive cracking of the infill panel. If an arching mechanism can be developed, the maximum lateral storey drift ratio of the slenderest panel permitted (slenderness equal to 20) is 2.8%, which is very near to the drift limit of 3% given for the Life Safety limit. Thus, all the infill panels that can develop arching mechanisms can meet this required Performance Level, since their strength will be enough to resist the inertia forces. FEMA-356 [55] recommends three different limit states, namely: (1) Immediate Occupancy Structural Performance Level: OOP drift ≤ 2%; (2) Life Safety Structural Performance Level: OOP drift ≤ 3%; (c) Collapse Prevention Structural Performance Level: OOP drift ≤ 5%. It is also indicated that if the surrounding frame is shown to remain stable following the loss of an infill panel, the remaining panels shall not be subjected to limits for the Collapse Prevention Structural Performance Level. Acceptable deformations of existing and new walls shall be assumed to be the same. Similar recommendations are provided by [56]. To conclude, some recommendations are provided by some authors based on experimental evidence. For example [4], propose a maximum ultimate OOP displacement equal to 0.8 times the panel thickness which corresponds to the activation point of the panel OOP instability. Kadysiewski et al. [57] and [12] suggested that the ultimate OOP displacement occurs five times the displacement corresponding to the maximum peak OOP strength. Based on the analysis of the test results available in this database, it was impossible to achieve solid conclusions regarding this parameter.

2.11 Final Considerations

39

Concerning the results obtained by the as−built panels, they are also characterized by some scatter due to the several variables involved as discussed in Sect. 3.4.

2.11 Final Considerations This chapter built a database of OOP experimental tests based on a systematic review methodology. The main variables and aspects that distinguish each testing campaign were summarised. In this systematic review methodology, the selection of the studies was filtered according to the scope of the present research. The complexity inherent to the infill panels’ OOP behaviour is reflected by the high number of parameters involved throughout the tests, such as the panel geometric properties, type of masonry units, boundary conditions, openings, loading protocol and previous damage due to IP test. This systematic review provides new findings that support the characterization of the OOP seismic behaviour of infill masonry walls. The major findings achieved from the global overview of the different testing campaigns are: – The most popular approach to applying the OOP loading is the use of airbags; – 12% of the database are combined IP−OOP tests; – Only 17% of the database cases are related to OOP tests combined the gravity load. Concerning this, some the authors point out the complexity inherent to the test setup as a limitation; – Large variations were found regarding the geometric characteristics of the panels tested available in the database. For example, it was observed that the panels’ slenderness and aspect ratio variation is around 10 to 56 and 0.4 to 1.6, respectively. The major observations that can be drawn from the as−built specimens available in the database are: – The maximum OOP strength capacity depends on the panel slenderness, namely higher maximum OOP strength loads were obtained by panels with lower slenderness; – The maximum OOP strength decreases for panels with larger aspect ratio; – Cracking patterns depend on the panel aspect ratio; – The infill panels follow a linear elastic behaviour up to the formation of the first crack. After that, the subsequent behaviour is nonlinear. This nonlinearity is related to the presence of new cracks and the propagation of existing ones up to the formation of the final cracking pattern; – The maximum OOP deformation capacity was observed in panels made with solid bricks. The results of panels made with solid bricks were included in the database information to increase the amount of data. However, the behaviour of infill panels made with solid bricks and hollow bricks is different thus the analysis and comparison between those results should be performed carefully;

40

2 Out−of−Plane Behaviour of Masonry Infill Walls Experimental Tests: …

– The infill panel properties play an important role in their OOP response, namely it was observed that larger compressive strength and flexural strength (both parallel and/or perpendicular to the horizontal bed joints) increased the maximum OOP strength capacity. An increment around 0.5 MPa of the flexural strength parallel to the horizontal bed joints increased the panel OOP maximum strength by five times. Finally, the major findings achieved from the study of the IP−OOP tests are: – The IP−OOP behaviour interaction is still an open issue due to the reduced amount of data available in the literature; – Based in the data, empirical relationships were proposed to predict the infill panel cracking stiffness, secant cracking stiffness, maximum strength and secant stiffness when subjected to IP−OOP loadings with reasonable accuracy; – It was observed that the previous damage due to IP prior loading demands reduce the panel OOP initial stiffness and strength and potentiate fragile failure, leading to unexpected OOP expulsions. The modification of the boundary conditions justifies this since the panel detachment from the envelope frame occurs when subjected to IP loadings. To complement the present study and global findings, an additional number of tests should be included in the database.

References 1. Moghaddam H, Dowling P, Ambraseys N (1988). Shaking table study of brick masonry infilled frames subjected to seismic actions. In: World Conference on Earthquake Engineering − 9WCEE. Tokyo, Japan. 2. Dawe J, Seah C (1989) Behaviour of masonry infilled steel frames. Can J Civ Eng 16:865–876 3. Dawe J, Seah C (1989) Out−of−plane resistance of concrete masonry infilled panels. Can J Civ Eng 16(6):854–864 4. Angel R, Abrams D, Shapiro D, Uzarski J, Webster M (1994) Behavior of reinforced concrete frames, with masonry infills. Civil Engineering Studies, Reserach Series No. 589, UILU−ENG, Department of Civil Engineering, University of Ilinois, USA, pp 94–2005. 5. Beconcini M (1997) Sulla resistenza a forze orizzontali di pareti in elementi forati in laterizio. Construire in laterizio 55:60–69 6. Flanagan R, Bennett R (1999) Bidirectional behaviour of structural clay tile infilled frames. J Struct Eng 125(3):236–244 7. Calvi G, Bolognini D (2001) Seismic response of reinforced concrete frames infilled with weakly reinforced masonry panels. J Earthquake Eng 5(2):153–185 8. Pereira P, Pereira M, Ferreira J, Lourenço P (2012) Behavior of masonry infill panels in RC frames subjected to in plane and out of plane loads. In: 7th Conference on on Analytical Models and New Concepts in Concrete and Masonry Structure. Cracow, Poland. 9. Varela−Rivera J, Polanco−May M, Fernandez−Baqueiro L, Moreno E (2012) Confined masonry walls subjected to combined axial loads and out−of−plane uniform pressures. Can J Civ Eng 39:439–447 10. Hak S, Morandi P, Magenes G (2014) Out−of−plane experimental response of strong masonry infills. In: Second European Conference on Earthquake Engineering and Seismology − 2ECEES. Turkey.

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11. da Porto F, Guidi G, Verlato N, Modena C (2015) Effectiveness of plasters and textile reinforced mortars for strengthening clay masonry infill walls subjected to combined in−plane/out−of−plane actions/wirksamkeit von putz und textilbewehrtem mörtel bei der verstärkung von ausfachungswänden aus ziegelmauerwerk, die kombinierter scheiben−und plattenbeanspruchung ausgesetzt sind. Mauerwerk 19(5):334–354 12. Furtado A, Costa C, Arêde A, Rodrigues (2016). “Geometric characterisation of Portuguese RC buildings with masonry infill walls.” Eur J Environ Civ Eng. 1–16. 13. Mosoarca M, Petrus C, Stoian V, Anastasiadis A (2016) Behaviour of masonry infills subjected to out of plane plane seismic actions part 2: experimental testing. In: International Brick and Block Masonry Conference − IB2MAC. Padua, Italy. 14. Silva L, Vasconcelos G, Lourenço P, Akhoundi F (2016) Experimental evaluation of a constructive system for earthquake resisting masonry infill walls. In: Brick and Block Masonry Conference (IB2MAC). Padua, Italy. 15. Liauw T, Kwan K (1992) Experimental study of shear wall and infilled frame on shake−table. In: World Conference on Earthquake Engineering − 10WCEE. Madrid, Spain. 16. Klingner R, Rubiano N, Bashandy T, Sweeney S, (1996) Evaluation and analytical verification of shaking table data from infilled frames. In: World Conference on Earthquake Engineering − 11WCEE. Acapulco, Mexico. 17. Fardis M, Bousias S, Franchioni G, Panagiotakos T (1999) Seismic response and design of RC structures with plan−eccentric masonry infills. Earthquake Eng Struct Dynam 28:173–191 18. Zarnic R, Gostic S, Crewe A, Taylor A (2001) Shaking table tests of 1:4 reduced scale models of masonry infilled reinforced concrete frame buildings. Earthquake Eng Struct Dynam 30:819– 834 19. Corte G, Fiorinho L, Mazzolani F (2008) Lateral−loading tests on a real RC building including masonry infill panels with and without FRP strengthening. Journal of Materials in Civil Enginering 20(6):419–431 20. Komaraneni S, Rai D, Eeri M, Singhal V (2011) Seismic behavior of framed masonry panels with prior damage when subjected to out−of−plane loading. Earthq Spectra 27(4):1077–1103 21. Stavridis A, Koutromanos I, Shing P (2012) Shake−table tests of a three−story reinforced concrete frame with maosnry walls. Earthquake Eng Struct Dynam 41:1089–1108 22. Tondelli M, Beyer K, DeJong M (2016) “Influence of boundary conditions on the out−of−plane response brick masonry walls in buildings with RC slabs.” Earthq Eng & Struct Dynamics 23. Augenti N, Parisi F, Acconcia E (2012) MADA: online experimental database for mechanical modelling of existing masonry assemblages. In: Fifteenth World Conference on Earthquake Engineering − 15WCEE, Lisbon, Portugal. 24. Kalman Šipoš T, Sigmund V, Hadzima−Nyarko M (2013) Earthquake performance of infilled frames using neural networks and experimental database. Eng Struct 51:113–127 25. Higgins J, Green S (2011) “Cochrane handbook for systematic reviews of interventions, Vol. 4.” John Wiley & Sons 26. Cadório I, Lousada M, Martins P, Figueired D (2017) Generalization and maintenance of treatment gains in primary progressive aphasia (PPA): a systematic review. Int J Lang Commun Disord 52(5):543–560 27. Reuters T (2016) EndNote X8 for windows & mac, released 8 November 2016. 28. Frederiksen V (1992) Membrane effect in laterally loaded masonary walls. A second order phenomenon. In: 6th Canadian Masonry Symposium. University of Saskatchewan, Canada. 29. Tu Y, Chuang T, Liu P, Yang Y (2010) Out−of−plane shaking table tests on unreinforced masonry panels in RC frames. Eng Struct 32:3929–3935 30. Liu M, Cheng Y, Liu X (2011) Shaking table test on out−of−plane stability of infill masonry wall. Transactions of Tianjin University 17(2):125 31. Varela−Rivera JL, Navarrete−Macias D, Fernandez−Baqueiro LE, Moreno EI (2011) Out−of−plane behaviour of confined masonry walls. Eng Struct 33(5):1734–1741 32. Guidi G, da Porto F, Benetta M, Verlato N, Modena C (2013) Comportamento sperimentale nel piano e fuori piano di tamponamenti in muratura armata e rinforzata. In: XV Convegno Nazionale ANIDIS − “L’INGEGNERIA SISMICA IN ITALIA”. Padova, Italy.

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33. Singhal V, Rai DC (2014) “Role of toothing on in−plane and out−of−plane behavior of confined masonry walls.” J Struct Eng 140(9):04014053. 34. Singhal V, Rai DC (2014) “Suitability of half−scale burnt clay bricks for shake table tests on masonry walls.” J Mater Civ Eng 26(4): 644–657. 35. Preti M, Migliorati L, Giuriani E (2015) Experimental testing of engineered masonry infill walls for post−earthquake structural damage control. Bull Earthq Eng 13(7):2029–2049 36. Moreno−Herrera J, Varela−Rivera J, Fernandez−Baqueiro L (2016) Out−of−plane design procedure for confined masonry walls. J Struct Eng 142(2):04015126 37. Akhoundi F, Lourenço P, Vasconcelos G (2016) Numerically based proposals for the stiffness and strength of masonry infills with openings in reinforced concrete frames. Earthquake Eng Struct Dynam 45(6):869–891 38. Akhoundi F, Vasconcelos G, Lourenço P, Silva B (2016) Out−of−plane response of masonry infilled RC frames: effect of workamnship and opening. In: International Brick and Block Masonry Conference − IB2MAC. Padua, Italy. 39. Silva L (2017) Experimental and numerical study of new systems for earthquake resistant masonry enclosures in reinforced concrete buildings, Research Thesis Project, University of Minho. 40. Di Domenico M, Ricci P, Verderame GM (2018) “Experimental assessment of the influence of boundary conditions on the out−of−plane response of unreinforced masonry infill walls.” J Earthq Eng. 1–39. 41. Akhoundi F, Vasconcelos G, Lourenço P (2018) Experimental out−of−plane behavior of brick masonry infilled frames. Int J Arch Herit 14(2):221–237 42. Di Domenico M, Ricci P, Verderame GM (2019) Experimental assessment of the out−of−plane strength of URM infill walls with different slenderness and boundary conditions. Bull Earthq Eng 17(7): 3959–3993. 43. Di Domenico M, Ricci P, Verderame GM (2019) Predicting the out−of−plane seismic strength of unreinforced masonry infill walls. J Earthq Eng: 1–38. 44. De Risi MT, Di Domenico M, Ricci P, Verderame GM, Manfredi G (2019) Experimental investigation on the influence of the aspect ratio on the in−plane/out−of−plane interaction for masonry infills in RC frames. Eng Struct 189:523–540 45. Furtado A, Rodrigues H, Melo J, Arêde A, Varum H (2019) Experimental assessment of strengthening strategy to improve the masonry infillsout−of−plane behaviour through textile reinforced mortar. In: COMPDYN 2019−7th ECCOMAS Thematic Conference on Computational Methods in Structural Dynamics and Earthquake Engineering, Crete, Greece. 46. Mosallam AS (2007) Out−of−plane flexural behavior of unreinforced red brick walls strengthened with FRP composites. Compos: Part B 38: 16. 47. Griffith M, Vaculik J (2007) Out−of−plane flexural strength of unreinforced clay brick masonry walls. TMS J 25(1):53–68 48. Griffith M, Vaculik J, Lam N, Wilson J, Lumantarna E (2007) Cyclic testing of unreinforced masonry walls in two−way bending. Earthquake Eng Struct Dynam 36:801–821 49. Ferreira T, Costa AA, Arêde A, Varum H, Costa A (2016) In situ Out−of−plane cyclic testing of original and strengthened traditional stone masonry walls using airbags. J Earthquake Eng 20(5):749–772 50. Lunn DS, Rizkalla SH (2011) Strengthening of infill masonry walls with FRP materials. J Compos Constr 15(2):206–214 51. Preti M, Bettini N, Plizzari G (2012) Infill walls with sliding joints to limit infill−frame seismic interaction: large−scale experimental test. J Earthquake Eng 16(1):125–141 52. Ricci P, Di Domenico M, Verderame GM (2018) Empirical−based out−of−plane URM infill wall model accounting for the interaction with in−plane demand. Earthq Eng & Struct Dyn 47(3): 802–827. 53. Ricci P, Di Domenico M, Verderame GM (2018) “Experimental investigation of the influence of slenderness ratio and of the in−plane/out−of−plane interaction on the out−of−plane strength of URM infill walls.” Constr Build Mater 191: 507–522.

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54. Fema−274 (1997) NEHRP commentary on the guidelines for the seismic rehabilitation of buildings: FEMA−274. Federal Emergency Management Agency, Washington (DC), Applied Technology Council, Washington, USA. 55. Fema−356 (2000) Pre−standard and commentary for the seismic rehabilitation of buildings, Federal Emergency Management Agency, Washington, DC. 56. ASCE (2013) Seismic evaluation and retrofit of existing buildings−ASCE41−13. American Society of Civil Engineers, Reston, Virginia 57. Kadysiewski S, Mosalam KM (2009) Modeling of unreinforced masonry infill walls considering in−plane and out−of−plane interaction, Pacific Earthquake Engineering Research Center, PEER 2008/102.

Chapter 3

Strengthening Strategies to Improve the Seismic Behaviour of Infill Masonry Walls

3.1 Introduction The present chapter aims to present retrofit and strengthening strategies to improve the infill masonry walls’ seismic performance, with particular attention to preventing the OOP collapse. Retrofit and improvement of infill walls’ seismic behaviour is a complex subject since it cannot be disconnected from their effect on the overall building response. It is paramount to assume this combined behaviour into consideration. The present chapter aims to present a literature review of the retrofit and strengthening techniques with revision of the major findings and observations made by each author. For each technique, a summary table will be provided containing the details regarding the design approach, major results and observations.

3.2 Literature Review of Retrofit and Strengthening Techniques The retrofit and strengthening of infill panels must be a priority nowadays both in new constructions and in existing constructions (e.g. rehabilitations). The OOP collapse vulnerability shown by the infill panels during the last few earthquakes, often mentioned in the previous chapters, reinforces the need for strengthening techniques to improve the infill walls’ performance in future events. Nowadays, the major concern during the design of a new structure is related to the structural members to ensure proper design of those with adequate capacity with the expected seismic demand. Few concerns are related to the “design” of the infill panels, considered by the codes as non-structural elements. The next generation of codes should demand the proper and adequate design of strengthening solutions to minimize their risk of collapse. Regarding the rehabilitation of structures, the strengthening and retrofit the infill panels distributed along the buildings’ façades should be a priority. © The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 A. F. Furtado, Seismic Vulnerability Assessment and Retrofitting Strategies for Masonry Infilled Frame Building, Springer Theses, https://doi.org/10.1007/978-3-031-20372-5_3

45

46

3 Strengthening Strategies to Improve the Seismic Behaviour of Infill … Infill walls

Approach 2

Approach 1

Sliding devices

Energy Dissipation Device

Gaps

New construction

Bed joint reinforcement

ECC

FRP

TRM

Existing construction

Fig. 3.1 Seismic retrofit and strengthening techniques of infill panels

According to the literature, two main approaches can be considered, as further described: (i) the disconnection of infills from the structural system (Approach 1); and (ii) the effective integration in the superstructure and strengthening of the panel (Approach 2). Both approaches can be adopted in the case of new constructions. On the other hand, in the case of existing constructions, the strengthening of the infill panels is more complex due to space limitations, lack of knowledge concerning the type of the infill panels’ masonry units and most important the lack of knowledge on how to proceed with the reinforcement. Figure 3.1 shows a schematic view of the available techniques to reinforce the infill panels. Approach 1 comprises three techniques: sliding devices, energy dissipation devices and the disconnection of the panel using gaps. Alternatively, Approach 2 is subdivided into four different strategies, namely the use of: fibre reinforced polymers (FRP), engineered cementitious composites (ECC), textile reinforced mortars (TRM) and bed joints’ reinforcement. Concerning the last technique, it is more usual in the case of new constructions.

3.2.1 Disconnection of the Infills from the Structural System 3.2.1.1

Sliding Devices

One of the most common techniques nowadays in the research works is the adoption of sliding devices to disconnect the infill walls partially from the structural system. This technique provides some capacity to dissipate energy to the wall and some deformation capacity (OOP or IP or combined IP + OOP), which is why those panels are called “engineered walls”. In 2011, Mohammadi et al. [1] conducted an experimental campaign to assess the performance of engineered infilled frames in two stages. One of the techniques used in Stage 1 was using an infill fuse, in which some sliding layers were included between the infill. In this technique, some elements, such as small parts of the columns or

3.2 Literature Review of Retrofit and Strengthening Techniques

47

horizontal layers in infills (called fuses), are supposed to yield or slide before infill cracking. Two 2/3 scaled, 3 m long and 2 m high single-story single-bay infilled steel frames having an IPE-140 standard shape were tested under cyclic lateral in-plane loading. The specimens were used to check the efficiency of the sliding devices to improve the panel ductility. The two tested specimens (SP3 and SF-SP3), were both composed of a 50-mm-thick concrete wall between two 100 mm masonry walls. The infill of the fused specimen (SF-SP3) included two sliding fuses. Each sliding fuse was composed of two (2.850 × 200 × 2 mm) steel plates with a thin layer of grease to reduce frictional resistance. Each plate of the fuse was welded to the reinforcement of the adjacent concrete wall. The authors found in previous experimental work that multileaf infill panels, composed of leaves of masonry and concrete materials, are acceptable to be used in engineered infilled frames, as they show higher ductility than single-leaf ones. Their strength can be adjusted by changing the layer thickness and material. The specimens were subjected to IP tests, where it was observed that the fused specimen evidenced higher deformation capacity and dissipated energy. The stiffness and strength degradation of the fused specimen (SF-SP3) was almost negligible, as opposed to the non-fused specimens (SP3), except for the last loading cycle with a drift amplitude higher than 7%, which rarely happens in earthquakes. For the fused specimen, the infill-frame interface cracked at the load of 31kN, after which the fuse started sliding. After that, neither infill cracking nor corner crushing was observed for drift demands up to 7.1% drift. The authors concluded that the use of sliding fuses had the advantage of increasing the deformation capacity and consequently the ductility of the infilled frame with the concrete wall, to meet current drift demands in earthquakes, normally being at a maximum of 2.5% and preventing the occurrence of damage/cracking during earthquake events. Despite the advantages of the sliding fuse, simple configuration of the applied sliding fuse had two main shortcomings: (1) increasing the vulnerability of potential shear failure in some parts of the columns and (2) creating a potential surface for outof-plane movement of the wall in the fuse area. The authors avoided the defects in the second specimens’ group through 30 mm chamfering located in the infills’ corners near the fuse and improved the fuse configuration to restrict transversal sliding. The fuse, herein referred to here as a frictional sliding fuse, was applied at midheight of the second group of specimens. It was composed of three steel plates: two of them (plates A and B) were fixed to each other by welding with a certain distance in between them, on which the third one (plate C) could slide. Plates B and C had some longitudinal slots and normal holes, respectively. Sliding between plates B and C is possible in the longitudinal direction but transversally restrained (to stabilize the wall in their OOP direction). Six high-strength N20 bolts connected these two plates, B and C, which were used to regulate the fuse for a certain sliding strength. In this testing campaign, the specimens EIF-1 to EIF-2 were subjected to IP and OOP tests. Those specimens have the same configuration; however, their fuses were set for different sliding strengths, 51kN and 73kN, respectively. During the IP test, similar behaviour was observed in both specimens: cracking of the infill-frame

48

3 Strengthening Strategies to Improve the Seismic Behaviour of Infill …

interface occurred initially, then 45° inclined cracks initiated in the infill near the shear connectors, followed by fuse sliding, which started in the 17th cycle at load and drift values of 80.3kN and 0.4%, respectively for EIF-1, and in the 30th cycle at load and drift values of 136.9kN and 0.5%, respectively for EIF-2. The secondary effect of the frame increases the sliding strength of the fuse by raising the normal load (based on the results of vertically installed strain gauges on infill material, immediately above the steel plate layers). As a general trend, the hysteretic loops were symmetric in both loading directions, except for a minor difference in the ultimate load capacities. The degradation of the stiffness and strength of the specimens with respect to repeating loads was almost negligible, by contrast with ordinary infill panels. In order to study the efficiency of this retrofitting strategy under OOP loadings, the specimen EIF-1 was driven back to its normal position, zero drift, after having failed by IP loading. Then, a manual jack applied the load perpendicularly at the infill centre. Throughout the test, the wall stayed stable and did not fall out. Preti et al. [2] conducted two experimental campaigns, focusing on developing a similar engineered solution comprising sliding joints to reduce the infill-frame interaction and ensure OOP stability. The authors validated the potential of horizontal partition joints (embedded in a few masonry mortar beds and acting as sliding joints) to ensure a ductile mechanism for the infill under IP loading; during the tests it was prevented the development of the typical diagonal strut mechanism. In the wall with partition joints, the continuity was broken by two couples of shaped steel folded plates having 2 mm thickness, placed at 1/3 and 2/3 of the panel height to reduce friction and force sliding along the partition joints. The shape of the steel folded plated was defined to increase the OOP infill stability using a shear key. The OOP failure occurred with a ductile mechanism for a load corresponding to an equivalent horizontal acceleration equal to 4 g. Such strength appeared not to be affected by several IP displacement cycles up to 2.5% drift, which was consistent with the marginal damage suffered by the infill for IP load, even at large deformation stages. More recently, Vailati et al. [3] developed a hybrid system comprising recycled plastic joints in place of mortar layers to join juxtaposed hollow clay blocks. Each joint is made of a planar surface with protruding teeth that fit into the holes of the masonry units. Based on the location of the joint (at the panel base, along the panel height, or at its top), the teeth are placed on one or both surfaces. The authors’ main objective was to provide a solution that allows dissipating energy when subjected to OOP seismic demands. Three full-scale specimens were tested at the LESE laboratory, using the test setup using airbags. The objective of the testing campaign was to assess the performance of the panel tested under different boundary conditions and surface finishing. Polyurethane was used to fill the last joint between the masonry units and the top beam. From the tests, the authors pointed out that the panels exhibited good behaviour under the OOP loadings by showing a high capacity to deform and dissipate energy. However, the authors mentioned that particular attention must be provided to the last top joint between the panel and the top beam, since the failure mechanism can be triggered by the shear sliding in the top of the wall.

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49

Energy Dissipation Devices

The use of energy dissipation devices consists of disconnecting the panel from the frame structure and the insertion of energy dissipation devices between them. The use of properly designed devices allows dissipating energy during an earthquake while being easily replaceable in a post-earthquake scenario. However, the design of this device needs adequate knowledge of the infill panel IP and OOP behaviour, which somehow needs to be interconnected. The first study of this type of solution was carried out by [4], in which a design criterium was formulated in terms of energy, providing an optimal balance of stiffness and energy dissipation to the structure through appropriate cladding connection. Aliaari et al. [5] tested a seismic infill wall isolator from the main envelope structure, designated by the authors as “SIWIS”. The solution consisted of using sub-frames attached to the structural frame and the infill wall is then constructed inside the sub-frame. The concept of SIWIS includes two vertical members and one horizontal member placed between an infill wall and the structural frame. The OOP stability of the panel was provided through the top sub-frame member. The authors stated that the location of SIWIS elements near the top of the wall was chosen because the frame will first contact the panel at that point under lateral drift and will tend to close the gap if it is not used any SIWIS elements. The authors simulated the effect of the SIWIS solution on the lateral capacity of an RC frame. From the analyses, the authors observed that this solution provides 10 times higher initial stiffness when compared to the bare frame configuration. The IP stiffness of the frame with SIWIS elements decreased by progressive phases until reaching the bare frame stiffness in the last phase after all SIWIS elements had failed. Later, Aliaari et al. [6] carried out an IP test of a two-bay, three-storey, steel frame with three different configurations, namely: (i) bare frame; (ii) infilled braced frame; and (iii) pinned frame equipped with a SIWIS device. The authors also tested a series of components with three different designs for the fuse element. From the tests, the authors pointed out that the frames with SIWIS elements showed an initial linear response followed by stiffness degradation due to progressive cracking of the SIWIS elements. Drops were observed concerning the lateral strength associated with the failure of each SIWIS element. The authors emphasize that this may be thought to impart sudden shocks to the building structure from a practical point of view.

3.2.1.3

Disconnection Using Gap

Seismically active countries such as New Zealand, Japan, Turkey and some states in USA adopted the practice of separating the infill panels from their RC frames by providing a gap between them. This strategy was based on the poor seismic performance of the infill panels in past events. Similarly, the seismic codes require that non-structural elements should not be damaged during earthquakes with low magnitude and do not affect the structural performance of the main structure in events with large magnitude. The separation between the panel and the frame became the

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most common practice [7], allowing the frame to deflect freely without mobilizing the wall contribution. According to the codes, the gaps between the infills and the structural frame must be sufficient and adequately calculated. Other demands such as acoustic insulation, weather tightness, fire protection and aesthetic qualities need to be addressed and compatibilized. Two essential features of a seismic separation gap are: (i) clear vertical gap between the panel and the columns (typically between 20 to 80 mm thickness); (ii) approximately 25 mm thick horizontal gap between the top of the panel and the top beam or slab. The gap under a beam or floor slab must be greater than the expected long-term deflection of that element, and also allow for the downwards bending deflection of a moment frame beam under seismic forces. These gaps allow the floor above an infill to move horizontally without the infill wall offering IP resistance [8]. This solution can be effective in terms of IP loading demand, but it considerably increases the panel OOP vulnerability. Some approaches are proposed to reduce this vulnerability increment, namely by: cantilevering the panel from its base; connecting the panel to the top upper floor to limit the OOP displacements (while maintaining an effective IP gap); or designing the infill wall to resist to the expected OOP loading demands and bending moments.

3.2.2 Infill Masonry Walls Retrofitting and Strengthening Techniques Large part of the RC Portuguese building stock was built before 1983, with improper seismic design and detailing and not providing any special attention to the infill walls and their interaction with the RC frames [9–11]. During the last few years, the solutions adopted for the RC building’s façade comprised different numbers of leaves (single panel or double-leaf), partial support width of the panels (for thermal insulation), different masonry units (clay or concrete with vertical or horizontal hollows), among others. However, the panel strengthening was always disregarded and not considered during the design of the building structure. The knowledge about strategies and techniques to improve the infill walls seismic behaviour have been an object of several studies and tests in the last years. However, in parallel to these advances, due to the concerns with thermal comfort, new bricks and techniques have also been developed for buildings’ façade walls to reduce cooling and heating losses. As a result of these innovations, new types of masonry units and construction technologies have been developed, being pushed by the market competition. The masonry industry improved the thermal properties of masonry units and developed new, faster and cheaper construction technologies [12]. External Thermal Insulation Composite Systems (ETICS) is now common in the external walls for energy saving purposes (Fig. 3.2). Distinct types of ties, generally from steel or plastic and having different shapes and geometry (dependent on the walls system)

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Fig. 3.2 Example of application of ETIC solution

are usually adopted [13]. However, it cannot be found in the literature any study regarding the effect of the ETICS on the infilled RC frames seismic performance. Based on the conception of this solution, apparently, no concerns related to the seismic performance were taken into account. Thus, it is important to develop and propose strengthening solutions to reduce the panels OOP collapse vulnerability. In the same way, as demonstrated in previous chapters, the existing panels built in the last decades are also vulnerable when subjected to earthquakes and need of retrofitting. The integration of the infill panels on the substructure and their strengthening allows the improvement of the seismic behaviour and the reduction of the panel OOP vulnerability can be achieved by adopting different strengthening techniques such as: fibre reinforced polymers (FRP), engineered cementitious composites (ECC), textile reinforced mortar (TRM) and bed joints reinforcement. The next sub-sections describe the recent developments for each strengthening technique, experimental results, application procedures and concerns related to the design methodology assumed by each author over the literature. At the end of each sub-section, a summary table containing the most relevant findings is presented.

3.2.2.1

Fibre Reinforced Polymers (FRP)

Carney et al. [14] tested two series of infill panels built with concrete blocks that were subjected to OOP loadings. A total amount of twelve walls with different strengthening configurations using FRP composite materials were tested. Two strengthening strategies using anchorages were adopted. The first method consisted of applying

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externally bonded glass FRP laminates. This strategy includes a primer and a glass fibre sheet to form the composite material. The authors stated that the glass fibre sheets are more economical and provide more compatible strength than the carbon fibres. The second method consisted of applying near surface mounted (NSM) glass FRP rods. These rods were attached to the wall using an epoxy-based grout. The anchorage of the laminates involved a groove in the boundary element, flush with the wall surface. To anchor the NSM rods to the boundary members, holes were drilled in the boundary in line with the head joints. The rods were glued with epoxy into the boundary at the same time they were applied to the wall. The specimens strengthened with anchorage produced a system able to support the double of the load of the reference specimen. Regarding the damages observed throughout the test, the failure mode was due to delamination, further propagated by tensile failure of the masonry near the location of the FRP. The authors concluded that the laminates increased the specimen capacity however, it was not the case for the specimen with NSM rods which reached lower strength due to the lack of bond and the limited bond transfer between the rod and the wall. In 2005, Hamid et al. [15] conducted an experimental investigation to study the IP behaviour of face shell mortar bedded infill wall assemblages retrofitted with FRP laminates. The testing campaign included compression strength tests parallel and perpendicular to the horizontal bed joints, diagonal tensile strength tests and shear strength tests. The FRP laminate reinforcement was selected according to an equivalent-stiffness-based approach, from which the laminate required was equal to the minimum steel reinforcement ratio of 0.2% (based on the gross cross-sectional area of the panel) according to the requirement of the Masonry Standards Joint Committee [16]. The Masonry Standards Joint Committee [16] requires a minimum steel ratio in high seismic prone regions to be distributed between the vertical and horizontal directions through the infill panels. Due to that, the authors determined the necessary thickness based on the premise that the FRP laminate’s stiffness would be at least equal or greater than the axial stiffness of the reinforcement in the walls. The required thickness of the FRP laminate was therefore calculated by direct scaling of the reinforcement area by the ratio of the elastic modulus of the steel and FRP material. From the testing campaign, the authors concluded that the laminates increased the load carrying capacity of the infill wall exhibiting shear failures along the mortar joints. The authors concluded that considering that the as-built specimens failed suddenly with total disintegration at the maximum peak load, the main contribution of the FRP in these cases was to improve the compression strength of the face shells by supplying the tensile strength required to stabilize the OOP buckling of the individual face shells, thus preventing failure after splitting of some parts of bricks. The FRP laminates resulted in a gradual prolonged failure under shear, and a stronger wall under compression with apparent post-peak strength. Almusallam et al. [17] carried out three IP tests on infilled RC frames with three different configurations, namely: (i) as-built; (ii) repaired specimen after prior IP test; and (iii) strengthened specimen. Hollow concrete blocks were used in the construction of the specimens. The specimens were repaired and strengthened on both sides

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of the panel, by using horizontal 300 mm wide sheet strips of a GFRP system, that were applied by using an epoxy resin by means of wet lay-up technique. These strips were used along the wall length parallel to the mortar lines. According to the authors, this retrofit strategy was selected for easy installation, cheaper solution and efficiency to resist shear stresses produced by horizontal loading demands. From the tests, the authors pointed out that the FRP improved behaviour the infill wall behaviour was shown by the increment of its strength and ductility. The deformation capacity of the panel increased up to 3 times [18] carried out two lateral loading quasi-static tests on a real masonry-infilled RC frame structure. The testing campaign started with the test of the original structure. This test found extensive damages in the masonry panels and RC elements and the occurrence of OOP collapse of almost all the walls parallel to the loading direction. After the test, the building was repaired, namely the façade panels were completely rebuilt and strengthened by FRP rods into the mortar bed joints. The test of the repaired structure revealed the efficiency of the near surface mounted FRP rods. The panels’ failure mode was modified from dominant diagonal cracking to prevailing shear sliding mode. It was not reported OOP collapse during the second test. Hrynyk et al. [19] performed several OOP quasi-static tests by applying a point load at the centre of the panels by using a hydraulic servo-actuator. In this testing campaign, the authors assessed the efficiency of modern materials in mitigating blast effects. The two strengthening materials under investigation were a glass fibrereinforced polymer grid and an elastomeric spray-on. From the tests, the authors concluded that when the strengthening material is applied in the entire face of the panel, the use of the GFRP grid increased significantly the load carrying capacity and prevented the material fragmentation upon failure. The authors observed that the strengthening material highly improved the deflection capacity of the specimens. This increment of deformation capacity led to significant improvements in terms of energy dissipation. Lunn et al. [20] conducted an extensive experimental campaign comprising 14 full-scale infilled RC frames specimens, four un-strengthened specimens and 10 strengthened specimens. Solid clay bricks were used to build the infill walls specimens. The strengthened specimens were reinforced with externally bonded glass fibre–reinforced polymer sheets applied in the external leaf of the panel. The authors adopted different thickness ratios considering only unidirectional (vertical or horizontal) directions. Regarding the anchorage systems of the strengthening material, the authors tested three strategies: GFRP overlap onto frame; no overlap onto frame; and shear restrain anchorage system. All the panels were subjected to pure OOP tests, in which distributed OOP loads with airbags were applied throughout the entire panel. The airbags were placed between the panel and a RC strong reaction wall. The RC frame was anchored to the reaction wall along 14 different points equally distributed through steel rebars. From the testing campaign, the authors concluded that the externally bonded solution was effective if proper anchorage of the FRP laminate is guaranteed. Overlapping the FRP reinforcement onto the RC frame was revealed to be very effective for double leaf specimens, however not so effective for single leaf specimens.

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The cracking pattern of the double leaf reference specimens experienced an initial horizontal crack, followed by a main vertical crack which dominated the panel behaviour. Strengthened specimens in which the FRP retrofit was terminated at the outer edge of the masonry with no overlap onto the RC frame failed by shear sliding mode, which was characterized by a large relative slip between the masonry and the RC frame. Specimens strengthened with the glass FRP sheets overlapped onto the RC frame failed by debonding mode characterized by delamination of the sheets, starting at the interface between the masonry and the RC frame. Finally, mechanically anchored specimens using the steel shear restraint anchorage system withstood over three times the design pressure without any visible strength degradation. However, the authors did not provide details regarding the meaning of the “design pressure”. Later, Valluzzi et al. [21] carried out an extensive experimental campaign of OOP flexural tests of masonry wallets made with HCHB to assess the impact of different strengthening techniques: application of organic (epoxy) matrices, i.e., FRP; Steel Reinforced Polymers (SRP), Natural FRPs, Carbon FRP, flax and hemp NFRP. Specimens were built with geometric dimensions equal to 390 × 1310 × 120 mm3 . The strengthening material was applied in one 50 mm wide central strip. Carbon, Basalt, Flax, Hemp and high density (HD) Steel sheets were applied as FRPs, i.e., with epoxy resin. Low Density (LD) Steel, Basalt and Glass meshes were applied with inorganic matrices, i.e., with cement mortar. The strengthening process of the specimens comprised the following steps, namely: – – – – –

Application of primer; Smoothing of the surface with a layer of putty; Application of a first layer of epoxy resin; Positioning of fibres; Use of a small paint roller to press the strip or a palette-knife, to allow proper impregnation of strands.

During the testing campaign, the authors observed that the failure mode was characterized by the fibres’ rupture in the case of Hemp FRP, shear failure in the masonry section for Basalt TRM and Carbon FRP solutions, combined ruptures (crushing and intermediate debonding) for Basalt FRP solutions and shear and debonding for SRP retrofit configurations. It was observed that all those failures occurred at the maximum peak load. The strengthened specimens with SRP and CFRP achieved the highest maximum loads. The remaining specimens reached under-average values, being the lowest strength reached by panels strengthened with natural FRP. Finally, the authors remarked that carbon, basalt and flax FRPs exhibited higher displacements at the peak load stage. Recently, Erol et al. [22] assessed the efficiency of CFRP strengthening technique to improve the lateral strength and stiffness of infilled RC frames. Mechanical characterization and IP tests were carried out in ½ scale specimens. Two different types of CFRP applications were adopted in this campaign, namely: (a) covering all the surface area of wall panel with CFRP fabrics with one layer in each diagonal direction; and (b) the other application was the strengthening of the wall panel with

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CFRP strips along the two diagonal directions. The CFRP strips were applied on both sides of the wall panels. The authors assumed variations concerning the CFRP strips widths, namely 300, 150 and 100 mm, for investigating a cost-effective solution. To prevent early CFRP strips’ debonding over the wall surface and benefiting from them until failure, CFRP anchors were used to bond the CFRP strips on both sides of the wall to each other in some specimens. CFRP fabrics were folded and impregnated with epoxy to form an anchor. From the tests, the authors concluded that the CFRP increased the lateral load capacity of the specimens from 1.2 to 1.8 times. The CFRP increased also the initial stiffness between 1.6 and 2.7 times. Regarding the damages observed, the corner crushing of the panels was prevented. The CFRP anchors that connected the diagonal CFRP strips on both sides of the wall were efficient until the failure of the frame with low-strength concrete and vulnerable beam-column joints’ detailing. Table 3.1 summarizes the major findings regarding the studies of strengthening of infill panels using FRP. This Table presents the information concerning the type of test carried out, type of masonry unit, provision of design approach by the author, specification if the strengthening strategy was designed for IP or OOP or both and the most relevant observations and findings drawn by each author. The works were grouped according to the type of test.

3.2.2.2

Engineered Cementitious Composites (ECC)

Engineered cementitious composites, commonly known as “ECC”, is a mortar-based composite reinforced with specially selected random fibres (usually steel, polymer, or carbon). The ECC can reach high levels of tensile strength and tensile strain capacity in the range between 3 and 7%. Thus, the ECC is considered a ductile material which has been studied for the application in different types of RC elements. The use of ECC to improve the infill panels’ seismic behaviour started in 2005, with the study performed by Kesner et al. [24], which focused on the application of ductile fibrereinforced mortar material (ECC) to improve the in-plane behaviour of infilled RC frames. The testing campaign observed that different levels of strength increment between 1.2 and 1.6 times and stiffness increment between 1.4 and 2.6 times can be achieved by varying the mix design of the ECC material and the amount of reinforcement in the mortar. The authors observed that the failure of panels with ECC started with visible cracks. Nonetheless, the reinforcing details of the strengthening material anchorage to the frame affected the failure mode. Later, Billington et al. [25] proposed a thin layer of sprayable ECC applicable to strengthen infilled RC frames subjected to IP loadings. The authors tested 2/3scale specimens, from which they concluded that the ECC improved the infill walls’ performance in terms of strength (about 17%) and deformation capacity (about 2.6 times). The authors also pointed out that, the retrofit details require special attention to bond the ECC layer to the infill panel and to connect the ECC to the frame since without an effective anchorage of the strengthening material to the frame, the performance of the strengthening solution can be affected.

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Table 3.1 Summary of testing campaigns of FRP strengthening of infill walls Author

Type of test

Masonry units

Design approach

Hamid et al. [15]

ME

SB

Yes

Erol et al. [22]

ME IP

HCHB

No

Almusallam et al. [17]

IP

SB

No

Corte et al. [18]

IP

SB

No

Carney et al. [14]

OOP

VHCB

No

x

Hrynyk et al. [19]

OOP

SB CSB

Yes

x

SB

Yes

x

Lunn et al. [20, OOP 23]

IP retrofit

OOP retrofit

√

√

√

√

√

Major observations and results – Increment of strength and OOP stability

x

– Increment of strength around 1.2 to 1.8 times; – Increment initial stiffness from 1.6 to 2.7 times

x

– Increment of strength and ductility – Deformation capacity increased up to 3 times

x

– Modification of the panel’s failure mode

√

√

√

– Increment of strength up to 2 times – Increment of the panel deformation capacity, energy dissipation and strength – Delamination of the GFRP sheets – Increase of maximum strength up to 3 times (continued)

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Table 3.1 (continued) Author

Type of test

Masonry units

Design approach

IP retrofit

Valluzzi et al. [21]

OOP

HCHB

No

x

OOP retrofit √

Major observations and results – Increase of maximum strength up to 2 times – Debonding of the FRP strips and masonry shear failure

ME—Mechanical characterization tests IP—In-plane tests OOP—Out-of-plane tests VHCB—Vertical hollow concrete block SB—Solid brick CSB—Concrete solid brick HCHB—Hollow clay horizontal brick

Kyriakides et al. [26] carried out an experimental program that comprised four quasi-static IP cyclic tests, on 1:5 scale infilled non-ductile RC frames, using solid clay bricks, three of them strengthened with ECC layer applied only in the external face of the panel. The fourth specimen (the reference one) was tested without retrofit. All the strengthened specimens used ECC material suitable for wet-mix shotcrete. The ECC mixture proportion was based on the design proposal by [28], and is summarized in Table 3.2. The Specimen EW-25 strengthening schematic layout was designed aiming at reducing the cracking and consequent failure of the infill panel. This specimen was strengthened with an ECC layer applied on one side of the panel, with a thickness equal to 12.7 mm, and reinforced with welded steel wire fabric amounting to 0.25% of the ECC layer cross-section, which corresponds to 127mm2 /m at full-scale surface. To improve the bond between the ECC and the masonry surface and limit the delamination while facilitating crack spread in the ECC in tension, wide strips of bonding agent were painted on the masonry surface before the application of the reinforcement and the ECC layer. The second strengthening proposal is intended to eliminate the failure of the masonry infill and delay the RC frame failure. The specimen EWBD-40 was strengthened with the ECC layer with bonded dowels and with 0.40% reinforcement in the ECC layer, corresponding to 203mm2 /m at full-scale surface. This strategy was similar to the one adopted in the specimen EW-25 in terms of the welded wire fabric attachment and the use of a bonding agent. It used ten shear dowels to connect the ECC layer to the top and bottom concrete beams, which provided a direct path for the lateral load being resisted by the ECC attached to the masonry wall. Holes were drilled with 76 mm depth at a roughly 25° angle (the angle depended on the reinforcement configuration) into the top and bottom RC beams and were used to accommodate the shear dowels, which were 127 mm segments of ø9.5-mm

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Table 3.2 ECC mixture proportion of sprayable (all ratios in volume except fibre volume fraction) adopted by Kyriakides et al. [27]

Material constituents

Ratio

Type I cement

0.95

Water

0.46

Silica and F-110

0.80

Class F fly ash

0.30

Hydroxypropylmethylcellulose

0.00075

High-range water-reducing admixture

0.0075

Calcium aluminate cement

0.05

fibre volume fraction

0.02

diameter reinforcing steel bars. The shear dowels were fastened in the concrete using epoxy resin. A symmetric response was observed in each loading direction, which was attributed to the ECC that, according to the authors, prevented any major crack formation in the panel. It was no observed crack in the ECC layer during the test, with the exception of a few ones in the infill wall corner. Shear failure of the columns at the top with almost no damage in the ECC evidenced a typical “strong” infill behaviour. No delamination of the ECC layer was observed. The third strengthening proposal intended to eliminate the failure of the masonry infill, to delay the RC frame failure and also to keep the lateral capacity of the frame structure for higher IP drift demands than those obtained with specimen EWBD40. This specimen, EWUD-40, was strengthened with ECC layer with unbounded dowels and with 0.40% reinforcement in the ECC layer. The strengthening proposal for the specimen EWUD-40 was similar to the one adopted in the panel EWBD-40, with only two exceptions, namely: (a) 8 shear dowels were used both at the top and bottom borders to connect the ECC layer with the RC top beam and base, respectively; and (b) the dowels used at the base were unbonded from the ECC. The aim of unbonding the dowels from the RC base beam was to allow the transfer of shear load and preventing the dowels from resisting to tensile loads due to their overturning. The shear dowels were applied in the RC base in a similar manner to the one adopted in specimen EWBD-40. However, duct tape and grease were applied after the dowel installation and before applying the ECC layer over and around the dowels. The as-built infill panel (UW) cracking pattern was characterized by bed joints sliding at the top combined with diagonal crack and cracks along the columns’ height. Specimen EW-25 exhibited undesirable horizontal shear failures at the top of the columns due to the strong infill not tied into the surrounding frame. Again, distributed cracking along the columns’ height was reported. No cracking was observed in the ECC layer or in the infill panel except the cracking along the top bed joint. No delamination between the ECC and the masonry was observed, indicating that strips of bonding efficiently improved the bond between the ECC and the panel surface. Distributed cracking occurred during the test of specimen EWBD-40 combined with

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cracking at the beam-column joints and some cracks concentrated at the base of the columns. The full bond developed between the ECC layer and the base dowels increased the lateral strength and stiffness of about 87% and 61%, respectively. The specimen EWBD-40 began to show rocking behaviour as the shear dowels at the base started to be pulled out. After the pull-out of the shear dowels, the specimen began to slide along the base. However, stable, ductile behaviour was observed, with specimen EWBD-40 maintaining 61% of its lateral capacity up to 3% IP drift. Regarding the effect of the strengthening of panel EW-25, it was observed that the initial stiffness increased about 20% and the maximum peak load about 18% compared to the reference specimen. The cracking pattern of the specimen EWUD-40 was similar to the one observed in specimen EWBD-40, however with larger shear sliding along the horizontal bed joints instead of the diagonal cracking. The specimen demonstrated ductile behaviour, maintaining 75% of its ultimate lateral strength up to 5.35% drift. The specimen exhibited 58% and 34% higher strength and stiffness, respectively. Concerning the effect of the strengthening in the energy dissipation capacity of the panels, the specimen EW-25 dissipated up to 50% more energy than the reference one UW. The retrofit schemes of EWBD-40 and EWUD-40 provided up to 2.5 times more energy dissipation than the un-retrofitted specimen UW. In the same year, Kyriakides et al. [27] carried out a research study on the effect of a thin layer of ECC in infill wallets, made with solid clay bricks, subjected to flexure strength tests. The variables herein studied were the use of anchors between the ECC layer and the panel to improve the ECC-panel bond and materialize the steel reinforcement ratios of the ECC layer in the form of welded wire fabric. From the tests, the authors observed that the ECC strengthening increased the flexural strength and stiffness by 45% and 53%, respectively. Multiple cracking distributed were observed in the strengthened panels. More recently, Barros [29] conducted a testing campaign of masonry wallets built with hollow clay horizontal bricks subjected to flexural strength tests parallel and perpendicular to the bed joints. The objective of the experimental campaign was to assess the efficiency of the ECC strengthening technique to improve the OOP capacity and to evaluate the effect of different ECC thicknesses. For that, 30 specimens were built with 600 × 600 mm and 150 mm thick geometric dimensions. For each type of test, 5 as-built specimens (Group R), 5 strengthened specimens with 10 mm ECC thick (Group A) and 5 strengthened specimens with 20 mm ECC thick (Group B), were tested. The geometric dimensions of these specimens are similar to those adopted in the present work, in the flexural strength tests, that will be detailed in the next Chapter. From the flexural strength tests parallel to the horizontal bed joints, it was observed that the failure mode of the as-built specimens, shown in Fig. 3.3a, was characterized by the detachment of the first or second row of bricks from the adjacent row which, according to the author, was controlled by the mortar-brick adhesion. Regarding the strengthened specimens, similar damages where observed in both groups namely the occurrence of shear failure due to the small geometry of the panel (small distance between the OOP loading application and OOP restrains). The remaining failures

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(a)

(b)

(c)

Fig. 3.3 Testing campaign carried out by [29]: damages observed after the flexural strength tests parallel to the horizontal bed joints a as-built specimens, b group A and c group B.

were characterized by the crushing of the bricks combined with one or two major horizontal cracks, shown in Fig. 3.3b and Fig. 3.3c and visible in the strengthened face of the panel. The tests showed that Group A and Group B specimens reached an average flexural strength equal to 0.43 MPa and 0.46 MPa, respectively. This means that the use of a double thickness of the ECC layer did not contribute significantly to the increment of strength, which was around 6%. The authors also pointed out that, due to the fragility of this type of masonry units, the ECC layer was too strong, which resulted in the amount of damage and crushing of the brick units. Regarding the comparison between the as-built specimens and the strengthened ones, the flexural strength was increased by about 5.38 times and 5.75 times for the Group A and Group B specimens, respectively. The authors compared also the OOP displacement capacity of each group corresponding to the maximum OOP peak load (df,oop,max ). The df,oop,max of the Group A and Group B specimens was 1.86 times and 2.04 times higher than the as-built specimens, respectively. The double ECC layer thickness of the Group B specimens contributed for a df,oop,max 9% higher. The flexural strength and OOP displacement corresponding to the maximum OOP load are plotted in Fig. 3.4a, b, respectively. Concerning the flexural strength tests perpendicular to the horizontal bed joints, the as-built specimens’ failure mode was characterized by a pure vertical cracking along the middle alignment of the wallet, as shown in Fig. 3.5a. Different failure modes were observed in the strengthened specimens. The most representative failure of the group A specimens was characterized by vertical cracking similarly to the asbuilt group, as shown in Fig. 3.5b. On the other hand, for group B specimens there was shear failure and masonry crushing (Fig. 3.5c, d). The author pointed out that the use of double thickness of the ECC layer led to higher levels of damage in the masonry and, as observed in the flexural strength tests parallel to the bed joints, did

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(a)

61

(b)

Fig. 3.4 Testing campaign carried out by [29]: flexural strength tests parallel to the horizontal bed joints a flexural strength and b df,oop,max

not explore the reinforcement material correctly. However, it is worth recalling that these results were, again, affected by the geometry of the specimens. From those flexural strength results, it was observed that Group A and Group B reached an average flexural strength equal to 1.31 MPa and 1.85 MPa, respectively. From this, it can be concluded that the double thickness of the ECC layer increased the OOP strength about 29%. Regarding the comparison between the as-built specimens and the retrofitted ones, the flexural strength was increased by about 2.52 times and 3.54 times for the Group A and Group B specimens, respectively. The df,oop,max of the Group A and Group B specimens was around 1.13 times and 1.88 times higher than the as-built specimens, respectively. The double thickness of the Group B specimens contributed for a df,oop,max 40% higher. The flexural strength results and the OOP displacement corresponding to the maximum OOP loading are plotted in Fig. 3.6a, b, respectively. Other studies were carried out concerning the efficiency of using ECC to improve the infill panels’ seismic behaviour [30–32]. Table 3.3 summarizes the most relevant studies available in the literature concerning the use of the ECC strengthening technique, where the major observations and results are listed. The table is organized according to the type of test and masonry unit type.

3.2.3 Reinforced Plaster Since 1980, the use of reinforced plaster started to be used in civil engineering. The reinforced plaster can be divided into two strategies: textile reinforced mortar (TRM) and steel meshes, for example. The TRM consists of using reinforcing meshes such as textile meshes with different continuous fibres, embedded in mortar which can have normal properties or high strength and/or ductility characteristics.

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(a)

(b)

(c)

(d)

Fig. 3.5 Testing campaign carried out by [29]: damages observed after the flexural strength tests parallel to the horizontal bed joints a as-built specimens, b group A, c Group B—detail of distributed damage along the specimen and d group c—masonry crushing

(a)

(b)

Fig. 3.6 Testing campaign carried out by [29]: flexural strength tests perpendicular to the horizontal bed joints a flexural strength and b df,oop,max

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Table 3.3 Summary of testing campaigns of ECC strengthening of infill walls. Author

Type of test

Masonry units

Design approach

IP retrofit

Kyriakides et al. [26, 27]

ME

SB

No

X

Dehghani et al. ME [31]

SB

No

Pourfalah et al. ME [32]

SB

No

x

HCHB

No

x

Barros [29]

ME

√

OOP retrofit √

–

X

√

√

Major observations and results – Increment of strength and stiffness of 45% and 53% under compression loadings – Flexural strength increased by approximately 3.5 times – Increase of the compression strength about 3–6 times and energy dissipation about 20–50 times – Increment of the shear strength about 1.5–2.8 times – Ductility increased 1.5 - 2 times – Increase of the flexural strength parallel and perpendicular to the horizontal bed joints 5.38 and 2.52 times, respectively; – Double thickness of the ECC layer increased the flexural strength 6% and 29%, respectively (continued)

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Table 3.3 (continued) Author

Type of test

Masonry units

Design approach

Kesner et al. [24]

IP

SB

No

Billington et al. IP [25]

SB

No

Kyriakides et al. [26, 27]

IP

SB

No

Dehghani et al. IP [31]

SB

No

Shing et al. [30]

SB

No

ST

ME—Mechanical characterization tests IP—In-plane tests ST—Shake table tests SB—Solid brick HCHB—Hollow clay horizontal brick

IP retrofit √

√

√

√

√

OOP retrofit

Major observations and results

X

– Increment of stiffness from 1.4 to 2.6 times; – Increment of strength from 1.2 to 1.6 times

X

– 2.6 times increase in ductility – 17% increase in strength

X

– Increase of the IP strength from 1.27 to 1.9 times – Increment of the energy dissipation capacity from 1.15 to 2 times

x

– The lateral strength and cumulative dissipated energy increased 2.57 and 2.74 times, respectively

x

– Increment of both the stiffness and strength

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The most basic application is the fibre reinforced mortar, which consists in a mixture of mortar with a percentage of fibres randomly distributed within its composition. It is generally used as shotcrete, which became widespread for tunnel reinforcements. Some factors that can affect the efficiency of this strengthening solution’s efficiency are the fibre slenderness and length, as well as the size of aggregates that compose the mortar matrix, since they define the bonding properties and consequently the capacity to behave as a composite [33]. More complex solutions using the same kind of material imply the definition of a principal direction for the reinforcement (according to the designer’s interest) along which the fibres can develop their maximum strength capacity. The constructive solutions can be divided into two solutions, namely unidirectional and bidirectional reinforcement meshes. The first investigations concerning the use of TRM based strengthening solutions were related to the tensile properties of the fibre reinforced mortars and their application on RC structures, namely in RC beams (to strengthen both bending and shear), or jacketing and/or confining RC columns. The application as a method of strengthening infill walls is a relatively new concept, proved by the reduced amount of investigations with full application in infilled RC frames. Among the parameters that can affect the performance of this strengthening strategy in infill panels, there are some remarkable ones that must be considered during the design procedure, namely: • Characteristics of the mesh, such as mesh density, which depends on the quantity of fibres in each thread (defined by the mass of textile reinforced mortar) and the thread spacing, the tensile strength and the ductility. Different types of textile meshes are available in the market, using different materials such as PVC, GFRP, CFRP or steel ones. Some examples of GFRP textile meshes are shown in Fig. 3.7; • Mortar and textile mesh interaction/adhesion properties can affect the bonding between the panel and the reinforcement material. The use of materials with compatibility and good adhesion properties are convenient for a better performance of the TRM solution. Several suppliers, recommend the application of the

Fig. 3.7 Examples of textile meshes, for TRM based solutions

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(a)

(b)

Fig. 3.8 Strengthening of infill panels using TRM based solutions: mortar-textile mesh interaction/adhesion a lateral view of the application of the mesh in a fresh plaster; and b lateral view of the TRM composite after an OOP test.

mesh in the first layer of a fresh plaster to improve the adhesion of the mesh to the plaster. For example, Fig. 3.8a shows the lateral view of applying a GFRP mesh in a fresh layer of plaster. Figure 3.8b shows the aspect of the textile mesh-mortar composite after OOP test; • Connectors to anchor the textile mesh to the panel and prevent delamination during a cyclic event, thus improving the strengthening solution’s performance. Different types of connectors are available in the market with different geometric characteristics and materials. Figure 3.9a shows examples of plastic connectors and their application is shown in Fig. 3.9b; • The anchorage of the textile meshes to the envelope RC frame elements, which is one of the most important details of the TRM based solution, since it is the parameter that may prevent the textile mesh detachment and consequently OOP collapse of the panel. As described later, the anchorage detail can affect the global performance of the strengthening solution. The anchorage can be performed through different types of elements such as metallic connectors, L-shape GFRP connectors, plastic connectors, among others, as shown in Fig. 3.10; • The overlapping of the mesh along the panel-frame transition, a region where high stresses can be expected to develop (Fig. 3.11a). The other critical region which requires overlapping of the textile mesh, is the transition between adjacent layers of the textile meshes, as shown in Fig. 3.11b. As mentioned before, the same concept of TRM has been also tested, but with other mesh types made of different materials, such as steel. In this line, [34] repaired

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67

(b)

Fig. 3.9 Strengthening of infill panels using TRM based solutions: textile meshes-panel connectors a examples of plastic connectors; and b application of the plastic connector

(a)

(b)

Fig. 3.10 Strengthening of infill panels using TRM-based solutions: textile meshes-frame connectors a examples of plastic, L-shape GFRP and metallic connectors; and b application of an L-shape GFRP connector

three specimens that were previously subjected to OOP monotonic tests. Hollow clay horizontal bricks were used to build the infill walls. The repair method consisted of spraying a half-inch thick ferrocement plaster to one or both faces of the infill panel. A single sheet wire mesh was applied on each face. The wire mesh was embedded in the plaster. Multiple wire sheets were lap spliced to cover the entire infill surface area. The wire mesh was anchored to the damaged infill panel by using steel bolts.

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(a)

(b)

Fig. 3.11 Strengthening of infill panels using TRM based solutions: overlap of the textile mesh a frame-panel transition; and b between layers of the textile meshes

(a)

(b)

(c)

Fig. 3.12 Testing campaign carried out by [45]: detail from the detachment from the plastic connector: a lateral view, b detail of the connector; and c detachment of the plaster from the column.

A flexible PVC sleeve was placed over the bolts to provide relaxation to the system and to prevent premature cracking of the plaster. The authors did not report details regarding the anchorage of the strengthening material to the envelope frame elements. From the tests, the authors concluded that the OOP maximum peak load increased about 5 times; and the OOP strength of the repaired specimens was not affected by the level of the previous damage. The authors also stated that the repair methods were very efficient due to the panels’

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lower slenderness ratio and that the compressive strength of the repairing material was high. Calvi et al. [35] tested the combination of using bed joints reinforcement with external reinforced plaster based solution. The adopted mesh was a welded wire metallic one. The strengthening was applied in both panel faces, being the meshes connected by steel plates in the bed joints, located at the distance equal to 1 m width and 0.6 m high, and bent around the wires of the meshes. The design approach was not provided by the authors as well the material properties of the mesh. From the testing campaign, the authors observed that the strengthening of the panel improved significantly the panel response, namely by increasing the deformation capacity and by modifying the damage limit states for higher drift levels. It was observed that the OOP acceleration demand to trigger the collapse of the panel was around 2.5 g without prior IP damage and between 0.5 and 0.7 g with prior IP damage. With the strengthening material, the acceleration increased to 3.5 g without prior IP damage and to 1.6 g with prior IP drift. Pereira et al. [36] carried out a testing campaign comprising six specimens subjected to combined IP-OOP test sequence, in scaled specimens built with hollow clay horizontal bricks 150 mm thick. Six specimens were built with different characteristics, namely: (a) infill panel, without plaster (Wall_REF_01); (b) infill panel, with 10 mm plaster in each face, Wall_REF_02; (c) infill panel, with bed joints reinforcement (JAR); (d) infill panel, with reinforced plaster composed by steel mesh with ø1.05 mm, with a mesh size equal to 12.7 × 12.7 mm (“Armanet” type), that was locally nailed to the RC element through fixing bolts using a fulminant technology and fixed to the metallic mesh with a lock washer 2 mm thick (Wall_RAR); (e) and (f) specimens Wall_PDA_01 and Wall_PDA_02 were double-leaf panels, with plaster in the external leaf and 10 mm gypsum in the internal leaf. The RAR specimen reached 10–25% higher maximum strength when compared with the as-built specimens. The cracking pattern of the as-built specimen Wall_REF_01 is characterized by the sliding cracks and diagonal cracks combined with crushing of some brick units. The strengthened specimen Wall_RAR response was different, namely the use of the TRM prevented the wall expulsion and allowed to distribute the cracking throughout the panel. Guidi et al. [37] performed combined IP-OOP tests intending to characterize the OOP behaviour of infill walls made with different types of masonry units, with and without strengthening. Two specimens were as-built, and two were built with reinforcement of horizontal and vertical bed joints. The remaining two specimens were built with a plaster layer, one of them strengthened by a special quadriaxial mesh made with hybrid glass fibres casted in an extra fibre-reinforced plaster layer. The authors provided no details concerning the material properties of the textile mesh. The thicker specimens, even when strengthened, developed bending OOP failure that somehow limited the panel strength. The OOP strength of reinforced infill walls were higher than those of unreinforced walls, for higher IP prior drift. Strength decay due to the increase of the IP drift was lower in the strengthened panel (−6%) than the one reached by the unstrengthened panels (−23%). The OOP strength of

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the strengthened wall thin walls was 54% higher than that of the un-strengthened wall. The displacement reached at 20% strength decay was 35% higher due to the strengthening of the panel. Koutas et al. [38] conducted a study concerning the development and performance of a new textile-based anchorage system used to transfer the tensile forces in infilled RC models, to ensure the connection between the panel and the RC frame elements. The authors proposed two different approaches to connect the infill panels to the frames, namely: (a) If the thickness of the panel is equal to the concrete frame elements width, the TRM layers can be extended from the panel to the concrete surface up to a practical, feasible length on both sides of the infill. Suppose the element of the surrounding frame is a T shape beam. In that case, the TRM layers can be extended up to the slab and possibly anchored there. In contrast, if the concrete element is a rectangular column, the TRM layers can cover the two faces of the column or even be wrapped around. Moreover, if the adjacent frame bay is infilled and the boundary has the same conditions, the TRM layers can be extended to the adjacent infill; (b) If the thickness of the panel is lower than the concrete frame elements width: simple extension of the TRM layers from the masonry to the concrete surface is not feasible and the transfer of forces from concrete to masonry is achieved through the use of anchors. The experimental program aimed to characterize the behaviour of new textilebased anchors as tension elements between TRM strengthened panels and RC prisms. Two different types of textile meshes were used for the external bonded reinforcement of the specimen, namely: (a) a commercial textile mesh with an equal quantity of epoxy-coated glass fibres in two orthogonal directions, mesh size equal to 10 × 10 mm and mesh density equal to 610 g/m2 , tensile strength equal to 147kN/m, rupture strain equal to 2.3%, and elasticity modulus equal to 70GPa); (b) lighter commercial textile mesh made of elastomeric polymer-coated E-glass with equal quantity of fibres in two orthogonal directions, with mesh size equal to 25 × 25 mm and mesh density equal to 405 g/m2 , tensile strength equal to 115kN/m, rupture strain equal to 2.5%, elasticity modulus equal to 73GPa. An additional textile mesh was used in this testing campaign to develop the anchors, which was used an uncoated basalt fibre roving knitted in two orthogonal directions with an equal quantity of fibres in each one, with a mesh size equal to 25 × 25 mm, mesh density 192 g/m2 , tensile strength equal to 66kN/m, rupture strain equal to 3.15% and elasticity modulus equal to 89GPa. The application of textile meshes and the textile-anchors comprised the following steps, namely: • Drilling holes into the concrete prisms with a diameter equal to 12 mm and varying the depth according to the specimen design. Removing the dust from the holes and from the panel surface that received the TRM layers with high-pressure air;

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• Application of protection to the holes drilled to keep them dry and clean during the application of the TRM layers; • Humidification of the surfaces that received the mortar; • Application of a thin layer of mortar at the joint between the concrete prism and the brick wallet; • Placement of the wallet on the concrete prism while the masonry mortar at the joint was still fresh; • Application of mortar on both sides of the wallet and of the concrete prism; • Bonding of the textile mesh by hand pressure; • Application of mortar at the location of the fanned part of the anchor; • Filling of the holes with low viscosity epoxy resin; • Local impregnation of the dry fibres of anchors with epoxy adhesive; • Placement of anchors into the holes, bonding the fan over the first textile layer by hand pressure and application of mortar on the top of the fan area; • Application of mortar in between the layers (in case of two layers) while the first layer was in a fresh state and on the top of the last layer. The thickness of each layer was approximately 3 mm, thus resulting in a plaster thickness of approximately 6 mm on each side of the specimens. At the location of the fan, the total thickness increased 2 mm. From the mechanical characterization tests, the authors pointed out that, the increment of the fibres quantity in the anchors resulted in a non-proportional increase of the forces carried by them. Due to improved bonding conditions, the anchors placed between two layers of textiles were more effective, about 50%, than those placed on top of a single layer. The most representative failure mode of the anchors was the debonding through the mortar followed by the rupture of the fibres. In the sequence of the study developed by [38, 39] performed an experimental campaign composed by two, full-scale, three-storey infilled RC frames subjected to IP loadings. The authors tested the use of TRM-based solutions as externally bonded reinforcement in combination with special anchorage details, similar to the solution developed and tested by [38]. The retrofitting strategy was composed by: column strengthening to prevent shear failure,strengthening of the infill walls via twosided application of layers of TRM externally bonded on the faces of the infills, and provision of adequate anchorage of the TRM around its perimeter via textile-based anchors and bond length. The strengthening application was made in seven steps: (i) shear strengthening of the columns located in the first and second storey columns in the plastic hinge lengths; (ii) after infilling the frames with the masonry and appropriate hardening, each panel was strengthened independently. The strengthening process started with the application of the first layer of TRM on both faces of the panel and around the boundary frame; (iii) application of the first TRM layer on the face of masonry infills, top part of the textile mesh; (iv) application of the textile anchors and extra textile patches on the front and back side of the specimen, respectively; (v) application of a second TRM layer on the faces of the first storey infill panel, the bottom part of the textile; (vi) application of second TRM layer on the faces of first story masonry infill,

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top part of the textile, and (vii) wrapping of the overhanging textile parts around the column corner. The tests observed that the strengthened specimens reached better results in terms of the lateral strength and deformation capacity, about 56% and 52%, respectively. The retrofitted specimen dissipated 22.5% more energy than the control one, for the same loading history. Francesca [40] assessed the efficiency of TRM based solutions on strengthening infill walls, made with HCHB, that were subjected to combined IP-OOP tests. Different solutions were tested, namely: (i) special lime-based plaster with geopolymer binder; (ii) bidirectional composite meshes with inorganic matrices (textile reinforced mortars TRM); (iii) TRM improved by anchorage of the mesh to the RC frame. The specimen 3-GC-NR was built with plaster prepared with natural hydraulic lime combined with glass and steel textile fibre mesh. The specimen 4-GC-FN was built with plaster prepared with natural hydraulic lime and geo-polymer binder with basalt and steel fibres mesh anchored with steel ties. The specimen 6-BG-NR was built with lime-gypsum plaster with basalt and steel fibres mesh. The specimen 8BC-NR was built with natural hydraulic lime plaster with basalt and steel fibres mesh. The authors did not provide material properties of the textile meshes. Wallet specimens were rendered on both sides with plaster made of natural hydraulic lime and geo-polymer binder (type GC, test series HB-UR), and with the same plaster and a glass and steel fibre mesh (test series HB-NR). The third series was composed of vertical strips of masonry (1298 × 384 mm), for flexural failure in the direction parallel to the mortar bed joints. In this case (VB-NR), the specimens were rendered with type GC plaster and strengthened with the mesh. The failure modes observed during the flexural tests were characterized by the cracking and failure of the mesh due to exceedance of the tensile strength. From the results, the authors pointed that the OOP behaviour of the panels improved significantly with plasters made of natural hydraulic lime, with or without geopolymer binder, reaching around 2.7 times to 3.5 times higher OOP strength. The TRM solutions allowed controlled collapse of the panel. The TRM did not affected the IP stiffness and strength capacity of the frame, but modified the cracking pattern and reduced the panel damages. The use of anchorages to connect the strengthening mesh to the top beam did not modify the infilled frame response compared with the specimens without the anchors. However, it provides an extra contribution concerning the damage control in the panel and the reduction of the displacement corresponding to the peak load, thus allowing a controlled failure. Martins et al. [41] tested strengthening techniques based on TRM with innovative reinforcing meshes based on composite rods comprising an external polyester protection of a reinforcing core with distinct types of fibres (called by the author “BCR” solution). The idea was to provide residual strength and deformation capacity, which controls the damage and avoids brittle failure of the masonry units. Flexural strength tests perpendicular to the horizontal bed joints were carried out in 5 masonry wallets, made with hollow clay horizontal bricks 150 mm thick: (i) 3 reference specimens without reinforcement; (ii) 3 specimens with commercial

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73

carbon mesh (comm_carb). The properties of the mesh are: mesh size equal to 25 × 25 mm, unidirectional mesh, mesh density equal to 200 g/m2 , tensile strength equal to 93.6kN/m and ultimate strain equal to 1.75%; (iii) 3 specimens with commercial glass mesh (comm_glass). The properties of the mesh are: 25 × 25 mm mesh size, 225 g/m2 mesh density and tensile strength equal to 45kN/m and ultimate strain equal to 3%; (iv) 3 specimens with a BCR mesh with a reinforcing core composed of 3 yarns of carbon and a braided structure of type 1b8min, and (v) 3 specimens with a BCR mesh with a reinforcing core composed of 5 yarns of glass and a braided structure of type 1b8min. From the tests, the authors concluded that flexural cracking load increased between 7 and 34% for manufactured mesh with glass and carbon fibres, respectively. The authors also found that cracked stiffness was about 50% and 10% of the initial stiffness in the case of meshes made with carbon fibre and glass fibre, respectively. The bending strength is higher in the specimens strengthened with the manufactured meshes with BCR, compared with the equivalent commercial meshes. Concerning the ductility factor, it is higher in the specimens strengthened with the manufactured meshes than in those reinforced with the commercial meshes. Therefore, Baloevi´c et al. [42] studied the effects of plaster mortar and TRM strengthening on the IP behaviour of 1/3 scale, one storey, one bay infilled steel frames built with cellular lightweight concrete blocks. Three variants were considered for this study, namely: (a) unreinforced infilled RC frame, (b) retrofitted infilled RC frame with pre-mixed plaster on both sides; and (c) retrofitted infilled RC frame with pre-mixed plaster on both sides and reinforced with a PVC mesh (mesh size equal to 3.6 × 3.6 mm; mesh density 160 g/m2 ). The authors concluded that the TRM strengthening increased the ductility (no details were provided regarding the quantification of the ductility improvement) and reduced the masonry’s cracks width. The ultimate bearing capacity was not affected by the strengthening. Recently, Akhoundi et al. [44] tested two strengthening TRM based solutions using two different textile meshes (glass fibre commercial and one developed by the authors in [41]. The properties of the commercial glass mesh are: tensile strength equal to 45kN/m, mesh density equal to 225 g/m2 , mesh size 25 × 25 mm, ultimate elongation equal to 1.8%. The properties of the textile mesh developed by the authors are: tensile strength and stiffness equal to 55kN/m and 154kN/m, density equal to 207 g/m2 and mesh size equal to 25 × 25 mm. Both textile meshes are bidirectional. Scaled infilled RC frames, made with hollow clay horizontal bricks were subjected to IP loadings. The authors’ main conclusion was that the TRM technique improved the IP response of infilled frame IP strength about 1.65 times and 1.70 times and the energy dissipation capacity from 2 to 3 times and works as damage concentrator. De Risi et al. [45] conducted a testing campaign comprising four full-scale specimens’ subjected quasi-static OOP tests using pneumatic actuators. One of them was tested in its as-built condition. The remaining three used different anchorages to connect the textile mesh to the surrounding frame (plastic connectors and L-shape GFRP connectors) and using normal or ductile mortar for the plaster. The authors used a GFRP textile mesh with a tensile strength equal to 40.0kN/m, and a mesh grid equal to 16.7 × 16.7 mm, supplied by Fassa Bortolo. The authors reported

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Fig. 3.13 Testing campaign carried out by [45]: detail from shear cutting

poor performance of the plastic connector as an anchorage element to the RC frametextile mesh even using a ductile mortar. The same problem was observed when using L-Shape GFRP connectors. The authors also pointed out that the shear cut of the textile mesh was observed in many cases due to the stress concentrations and by sliding of the mesh (see Fig. 3.13). This cutting and sliding phenomena contributed to the poor inefficiency of the adopted TRM solution. Globally, the authors found a good performance of the TRM solutions since they increased cracking and maximum strength around 1.77 to 2.22 times, respectively. The energy dissipation of each strengthened panel was significantly higher than the as-built specimens. The increment provided by the strengthening solutions for the same OOP displacement level was around 1.93 times. Finally, some other authors, such as [46, 47] and [48] carried out flexural strength tests to characterize the effect of the TRM technique in the improvement of the infill panels OOP capacity. Table 3.4 summarizes a list of the studies available in the literature concerning the use of TRM-based strengthening solutions. The major observations and results found by the authors are included.

3.2.4 Bed Joints Reinforcement The use of bed joints reinforcement is a common practice in some countries and consists of the use of reinforcement in the horizontal and/or vertical bed joints (see Fig. 3.14). This practice, as mentioned before is most common in new constructions. The implications of using this approach in existing constructions implies labour work which could increase the costs associated, thus making it impracticable. Dawe et al. [49, 50] tested the benefits of using reinforcement in the panel bed joints to improve the OOP behaviour. For this, specimen 7a incorporated bed joint

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Table 3.4 Summary of testing campaigns of TRM strengthening of infill walls Author

Type of test

Masonry units

Design approach

IP retrofit

Koutas et al. [38]

ME

HCHB

No

x

Martins et al. [41]

ME

HCHB

No

x

Baloevi´c et al. [42]

ME

CLCB

No

Ismail et al. [46]

ME

VHCB

Yes

Kariou et al. [47]

ME

SB

No

Shabdin et al. [48]

ME

SB

No

√

√

x

√

OOP retrofit √

√

Major observations and results – Anchors provided 2–3 times higher strength – Increase of the cracking load about 7% to 34% – Increase of the maximum flexure strength about 3 to 5 times

x

– PVC mesh reinforcement in the plaster does not contribute significantly to the load-bearing capacity of the structure

x

– Modification of the failure mode – Increase of the shear strength from 68 to 258%

√

x

– The maximum strength increased at least 1.8 times (with coated glass fibre textile TRM) – Improvement of the diagonal tensile strength and deformation capacity – The strengthened wall reached higher shear stress and pseudo-ductility, dissipated more energy – Connectors must be designed for OOP lateral force to ensure that the masonry wall and TRM reinforcement layer work together properly (continued)

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Table 3.4 (continued) Author

Type of test

Masonry units

Design approach

IP retrofit

Koutas et al. [39]

IP

HCHB

No

Akhoundi et al. [43, 44]

IP

HCHB

No

Calvi et al. [35]

IP-OOP

HCHB

No

Pereira et al. [36]

IP-OOP

HCHB

No

Guidi et al. [37] IP-OOP

HCHB

No

da Porto et al. [40]

IP-OOP

HCHB

No

x

Valluzzi et al. [21]

OOP

HCHB

No

x

√

√

OOP retrofit

Major observations and results

x

– Increase of the lateral about 56% increase, the deformation capacity 52% and the energy dissipation 22.5%

x

– Increase of the stiffness and lateral strength; – Post-peak behaviour is controlled

√

√

√

√

√

√

√ √

– Modification of the damage limit states – 4 times higher stiffness; – Higher maximum strength around 10–25% – Higher OOP strength for specimens damaged due to prior and higher IP drift – Increase of OOP strength around 30% – 2.7 times higher OOP strength – Increase of maximum strength 3 times (continued)

reinforcement at alternated courses and during the tests it was verified a considerable increase of the panel’s ductility. The reinforced specimens reached higher strength corresponding to the first crack. No significant differences were observed between the specimens with and without bed joint reinforcement. The authors also reported the increment of strength at the occurrence of the first crack due to the bed joints reinforcement. Finally, Silva [51] carried out an experimental campaign comprising 8 OOP and IP-OOP tests to assess the seismic behaviour two innovative solutions, that consisted of the use of properly designed masonry units combined with the introduction of reinforcement bars to improve their seismic capacity. Another objective was to improve

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Table 3.4 (continued) Author

Type of test

Masonry units

Design approach

IP retrofit

Angel et al. [34] OOP

HCHB

No

x

De Risi et al. [45]

HCHB

No

x

OOP

ME—Mechanical characterization tests IP—In-plane tests IP-OOP—Combined IP-OOP test sequences SB—Solid brick HCHB—Hollow clay horizontal brick VHCB—Vertical hollow concrete block CLCB—cellular lightweight concrete blocks

Fig. 3.14 Examples of bed joint reinforcement

OOP retrofit √

√

Major observations and results – OOP maximum strength increased 4 times; – Reduction of the post-peak OOP strength degradation – Increase of the maximum OOP strength between 1.77 and 2.22 times; – Increase the energy dissipation capacity about 1.93 times; – Different connectors were tested and if a proper connector is designed, the strengthening can be affected

78

3 Strengthening Strategies to Improve the Seismic Behaviour of Infill …

the thermal insulation characteristics of the panel. For this purpose, two systems were proposed by the authors, namely: The Uniko and the Termico systems. The Uniko system was designed for a situation of a single-leaf clay masonry enclosure to be taken for low to medium height residential/commercial RC frame structure and for zones of medium intensity seismic actions. The Uniko System is composed of VHCB masonry units. This masonry unit uses interlocking along the length. This system allows to dissipate high energy levels due to the interlocking joints, particularly the sliding between the masonry units. The masonry units were placed aligned with the panel length in the vertical direction and thus, with continuous vertical joints. Additionally, steel rebars were introduced in the face of the masonry units connected to the top and bottom RC beams/slabs. The second proposal included infill walls with dry vertical joints and mortared horizontal bed joints. The “Termico” system is characterized by maintaining the infill rigidly attached to the frame, using internal reinforcement along the horizontal bed joints and connectors between the infill and frame. The testing campaign observed an interesting improvement of the panels’ capacity in terms of both strength and deformation capacity. Globally, the Uniko system results were: (i) 40–50% higher OOP strength for undamaged and damaged panels; (ii) 20– 40% increment of the displacement at the OOP maximum strength. The Termico system obtained 19% maximum strength increment for the undamaged panel and 33% for a panel subjected to previous 0.54% drift. However, the specimen subjected to 1.05% IP drift reduced 20% of the maximum strength. The author provided no technical details to justify this result.

3.3 Final Considerations The present chapter aims presenting a literature review the existing seismic retrofit/strengthening techniques for infill walls. A detailed compilation and organized review based on their common characteristics and effects on the performance under seismic action were presented. Regarding the literature review, the following observations can be drawn: • Two different strategies can be adopted starting from the disconnection of the panel from the surrounding frame, energy dissipation, sliding devices, gaps, or effective strengthening of the panel based on different techniques such as ECC, TRM or FRP. Details regarding the retrofit approaches adopted by each author were presented. For each technique, a summary table including the most relevant information and results obtained were compiled; • Regarding strengthening the infill walls, all the techniques revealed interesting results, in terms of increasing the deformation capacity of the panels, preventing brittle failures. Depending on the technique, different results were observed in terms of strength, stiffness and energy dissipation improvements;

References

79

• From a technical point of view, all techniques are effective if the retrofit material is bonded or anchored to the panel and the surrounding frame. Different types of anchors were found in the literature, with different materials, with different application procedures, etc., but the retrofit strategy may be ineffective without accurate anchors design. Delamination and shear failure of the walls can suddenly occur without exploring all the potentialities of the strengthening material. The use of bed joints reinforcement is pointed as a solution with good efficiency since it provides deformation capacity to the panel. However, this strategy is related to as-built infill wall construction. Very few details are provided concerning the OOP performance of the strengthening solutions, besides discussing the maximum strength, energy dissipation and deformation capacity. Cost–benefit analysis would be important to assess the most adequate strengthening strategy according to the budget limitation. Most of the studies did not provide any design approach or guidelines that would be very important for structural designers to implement in civil engineering constructions. Different assumptions can be made for the retrofit design, in terms of pure prevention of OOP failure/collapse, optimizing the IP energy dissipation and reducing damages for lower PGA levels. The main challenge in the near future is to develop strengthening solutions which can improve the infill walls’ seismic behaviour and the global behaviour of the structure.

References 1. Mohammadi M, Akrami V, Mohammadi-Ghazi R (2011) Methods to improve infilled frame ductility. J Struct Eng 137(6):646–653 2. Preti M, Bettini N, Plizzari G (2012) Infill walls with sliding joints to limit infill-frame seismic interaction: large-scale experimental test. J Earthq Eng 16(1):125–141 3. Vailati M, Monti G, Di Gangi G (2018) Earthquake-safe and energy-efficient infill panels for modern buildings. In: Rupakhety R, Ólafsson S (eds) Earthquake engineering and structural dynamics in memory of ragnar sigbjörnsson: selected topics. Springer International Publishing, Cham, pp 233–261 4. Goodno B, Pinelli J, Craig J (1996). An optimal design approach for passive damping of building structures using architectural cladding. In: World conference on earthquake engineering— 11WCEE, Acapulco, Mexico 5. Aliaari M, Memari AM (2005) Analysis of masonry infilled steel frames with seismic isolator subframes. Eng Struct 27(4):487–500 6. Aliaari M, Memari AM (2007) Experimental evaluation of a sacrificial seismic fuse device for masonry infill walls. J Archit Eng 13(2):111–125 7. Charleson A (2008) Seismic design for architects. Elsevier Architectural Press 8. Okazaki T, Nakashima M, Suita K, Matusmiya T (2007) Interaction between cladding and structural frame observed in a full-scale steel building test. Earthq Eng Struct Dyn 36(1):35–53 9. Furtado A, Costa C, Arêde A, Rodrigues H (2016) Geometric characterisation of Portuguese RC buildings with masonry infill walls. Eur J Environ Civil Eng 1–16 10. Furtado A, Rodrigues H, Arêde A, Varum H (2016) Experimental evaluation of out-of-plane capacity of masonry infill walls. Eng Struct 111:48–63 11. Furtado A, Rodrigues H, Arêde A, Varum H (2016) Simplified macro-model for infill masonry walls considering the out-of-plane behaviour. Earthq Eng Struct Dyn 45(4):507–524

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12. Tomaževiˇc M, Lutman M, Bosiljkov V (2006) Robustness of hollow clay masonry units and seismic behaviour of masonry walls. Constr Build Mater 20(10):1028–1039 13. Collina A, Lignola G (2010) The external thermal insulation composite system (ETICS) more than comfort and energy saving. In: Third portuguese congress on construction mortars— 3PCCM 14. Carney P, Myers JJ (2003) Shear and flexural strengthening of masonry infill walls with FRP for extreme out-of-plane loading. In: Architectural engineering conference (AEI), Austin, Texas, EUA 15. Hamid AA, El-Dakhakhni WW, Hakam ZHR, Elgaaly M (2005) Behavior of composite unreinforced masonry-fiber-reinforced polymer wall assemblages under in-plane loading. J Compos Construct 9(1):73–83 16. MSJC (2004) Building code requirements for masonry structures (ACI 530–04/ASCE 5–04/ TMS 402–04) 17. Almusallam TH, Al-Salloum YA (2007) Behavior of FRP strengthened infill walls under inplane seismic loading. J Compos Construct 11(3):308–318 18. Corte G, Fiorinho L, Mazzolani F (2008) Lateral-loading tests on a real RC building including masonry infill panels with and without FRP strengthening. J Mater Civil Eng 20(6):419–431 19. Hrynyk TD, Myers JJ (2008) Out-of-plane behavior of URM arching walls with modern blast retrofits: experimental results and analytical model. J Struct Eng 134(10):1589–1597 20. Lunn DS, Rizkalla SH (2011) Strengthening of infill masonry walls with FRP materials. J Compos Constr 15(2):206–214 21. Valluzzi MR, da Porto F, Garbin E, Panizza M (2014) Out-of-plane behaviour of infill masonry panels strengthened with composite materials. Mater Struct 47(12):2131–2145 22. Erol G, Karadogan HF (2016) Seismic strengthening of infilled reinforced concrete frames by CFRP. Compos B Eng 91:473–491 23. Lunn DS, Rizkalla SH (2014) Design of FRP-strengthened infill-masonry walls subjected to out-of-plane loading. J Compos Constr 18(3):A4013002 24. Kesner K, Billington SL (2005) Investigation of infill panels made from engineered cementitious composites for seismic strengthening and retrofit. J Struct Eng 131(11):1712–1720 25. Billington S, Kyriakides M, Blackard B, Willam K, Stravidis A, Shing PB (2009) Evaluation of a sprayable, ductile cement-based composite for the seismic retrofit of unreinforced masonry infills. In: ATC and SEI conference on improving the seismic performance of existing buildings and other structures 26. Kyriakides MA, Billington SL (2014) Behavior of unreinforced masonry prisms and beams retrofitted with engineered cementitious composites. Mater Struct 47(9):1573–1587 27. Kyriakides MA, Billington SL (2014) Cyclic response of nonductile reinforced concrete frames with unreinforced masonry infills retrofitted with engineered cementitious composites. J Struct Eng 140(2):04013046 28. Yun Yong Kim H-JK, Victor CL (2003) Design of engineered cementitious composite suitable for wet-mixture shotcreting. Mater J 100(6) 29. Barros S (2017) Caracterização de argamassas com fibras de PVA para reforço de paredes de alvenaria MSc Thesis 30. Shing PB, Koutromanos I, Stavridis A (2013). Seismic performance of masonry-infilled RC frames with and without retrofit. J Earthq Tsunami 7(3) 31. Dehghani A, Fischer G, Nateghi F (2015) Strengthening masonry infill panels using engineered cementitious composites. Mater Struct 48(1):185–204 32. Pourfalah S, Suryanto B, Cotsovos DM (2018) Enhancing the out-of-plane performance of masonry walls using engineered cementitious composite. Compos B Eng 140:108–122 33. Gómez J (2012) Innovative retrofitting materials for brick masonry infill walls MSc Thesis, University of Minho 34. Angel R, Abrams D, Shapiro D, Uzarski J, Webster M (1994) “Behavior of reinforced concrete frames, with masonry infills, Civil Engineering Studies, Reserach Series No. 589, UILU-ENG, Department of Civil Engineering, University of Ilinois, USA, pp 94–2005

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35. Calvi G, Bolognini D (2001) Seismic response of reinforced concrete frames infilled with weakly reinforced masonry panels. J Earthq Eng 5(2):153–185 36. Pereira P, Pereira M, Ferreira J, Lourenço P (2012) Behavior of masonry infill panels in RC frames subjected to in plane and out of plane loads. In: 7th conference on on analytical models and new concepts in concrete and masonry structure cracow, Poland 37. Guidi G, da Porto F, Benetta M, Verlato N, Modena C (2013) Comportamento Sperimentale nel Piano e Fuori Piano di Tamponamenti in Muratura Armata e Rinforzata. XV Convegno Nazionale ANIDIS - “L’INGEGNERIA SISMICA IN ITALIA”. Padova, Italy 38. Koutas L, Pitytzogia A, Triantafillou TC, Bousias SN (2014) Strengthening of infilled reinforced concrete frames with trm: study on the development and testing of textile-based anchors. J Compos Constr 18(3):A4013015 39. Koutas L, Bousias SN, Triantafillou TC (2015) Seismic strengthening of masonry-infilled RC frames with TRM: experimental study. J Compos Constr 19(2):04014048 40. da Porto F, Guidi G, Verlato N, Modena C (2015) Effectiveness of plasters and textile reinforced mortars for strengthening clay masonry infill walls subjected to combined in-plane/outof-plane actions/Wirksamkeit von Putz und textilbewehrtem Mörtel bei der Verstärkung von Ausfachungswänden aus Ziegelmauerwerk, die kombinierter Scheiben- und Plattenbeanspruchung ausgesetzt sind. Mauerwerk 19(5):334–354 41. Martins A, Vasconcelos G, Fangueiro R, Cunha F (2015) Experimental assessment of an innovative strengthening material for brick masonry infills. Compos B Eng 80:328–342 42. Baloevi´c G, Radni´c J, Grgi´c N, Matešan D (2016) The application of a reinforced plaster mortar for seismic strengthening of masonry structures. Compos B Eng 93:190–202 43. Akhoundi F, Vasconcelos G, Lourenço P (2018) Experimental out-of-plane behavior of brick masonry infilled frames. Int J Archit Heritage 14(2):221–237 44. Akhoundi F, Vasconcelos G, Lourenço P, Silva LM, Cunha F, Fangueiro R (2018) In-plane behavior of cavity masonry infills and strengthening with textile reinforced mortar. Eng Struct 156:145–160 45. De Risi M, Furtado A, Rodrigues H, Melo J, Verderame G, Arêde A, Varum H, Manfredi G (2020) Experimental analysis of strengthening solutions for the out-of-plane collapse of masonry infills in RC structures through textile reinforced mortars (In Press). Eng Struct 46. Ismail N, El-Maaddawy T, Khattak N, Najmal A (2018) In-plane shear strength improvement of hollow concrete masonry panels using a fabric-reinforced cementitious matrix. J Compos Construct 22(2):04018004 47. Kariou FA, Triantafyllou SP, Bournas DA, Koutas LN (2018) Out-of-plane response of masonry walls strengthened using textile-mortar system. Constr Build Mater 165:769–781 48. Shabdin M, Zargaran M, Attari NKA (2018) Experimental diagonal tension (shear) test of UnReinforced Masonry (URM) walls strengthened with textile reinforced mortar (TRM). Constr Build Mater 164:704–715 49. Dawe J, Seah C (1989) Behaviour of masonry infilled steel frames. Can J Civil Eng 16:865–876 50. Dawe J, Seah C (1989) Out-of-plane resistance of concrete masonry infilled panels. Can J Civil Eng 16(6):854–864 51. Silva L (2017) Experimental and numerical study of new systems for earthquake resistant masonry enclosures in reinforced concrete buildings, Research Thesis Project, University of Minho

Chapter 4

Experimental Characterization of the As-Built Masonry Infill Components’ Properties

4.1 Framework Infill panels are widely used for partition purposes and to provide thermal and acoustic insulation to RC structures. Many types of masonry units were developed throughout the last decades to optimize those characteristics or to use innovative and green materials. The recent codes’ requirements regarding the thermal and acoustic insulation of buildings also motivated the development of new materials and new construction techniques of the buildings’ façades. The strategy adopted to build the buildings’ façades modified slightly over the years in Portugal. It started with the use of single leaf infill panels made with hollow clay horizontal brick (HCHB) units which were later replaced using double-leaf walls composed of HCHB units with different thicknesses. Examples of buildings built in Portugal, Spain and Italy with HCHB units along the façades are shown in Figs. 4.1 and 4.2. More recently, it became more commonly used single leaf infill panels made with blocks with special thermal and acoustic properties, as well as panels made with HCHB units combined with the application of an external thermal insulation composite system. The blocks herein mention to have special thermal and acoustic properties and are usually made of lightweight concrete or clay with vertical hollows. For this reason, it constitutes an alternative solution to the double-leaf solutions with light insulation in the intermediate air box and to the simple wall solutions with light insulation in the outside surface (ETICS). Due to this, this masonry unit presents an interesting breathing capacity, which avoids undesirable condensations inside the dwellings. Examples of buildings recently constructed with vertical hollow concrete or clay block units are shown in Fig. 4.2. A brief summary of the historic evolution of masonry units adopted in Portugal for the façades of RC structures can be found in Fig. 4.3, being important to mention that most of the existing buildings in Portugal were built with HCHB units.

© The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 A. F. Furtado, Seismic Vulnerability Assessment and Retrofitting Strategies for Masonry Infilled Frame Building, Springer Theses, https://doi.org/10.1007/978-3-031-20372-5_4

83

84

4 Experimental Characterization of the As-Built Masonry …

Aveiro

Seixal

a)

b)

Spain

Italy

c)

d)

Fig. 4.1 Examples of RC building structures built with HCHB units in the façades

The codes focus and requirements regarding functional characteristics are increasing, and it is notorious the continuous increment of demands with the aim of maximizing the comfort of buildings’ users. The contribution of masonry infill walls to the structural response of reinforced concrete buildings when subjected to an earthquake continues to be disregarded or ignored by some structural codes. However, several studies have developed simplified methods to consider the contribution of the infill masonry panels in the structural response. In fact, remarkable advances have been observed during the last decade in terms of development of numerical modelling frameworks, which are resorting to strut models’ concepts or detailed modelling approaches capable of simulating the infill panels’ seismic behaviour with good accuracy [1]. The strut models’ concept was originally developed to capacitate the numerical analyses of infilled frame structures. From its evolution, multi-strut models were developed to integrate the capabilities of the model to simulate the shear behaviour and tensile stresses within the contact length between wall and frame [2]. Numerical models have become increasingly complex, with some of them considering the infills’ stiffness and strength reduction under dynamic loadings or the adoption of other equivalent approaches to simulate the shear sliding failure of the infill walls. Recently, have emerged some developments regarding the strut models that

4.1 Framework

85

Loures

Lisboa

a)

b)

Aveiro

Lisboa

c)

d)

Fig. 4.2 Examples of RC building structures built with a, b Vertical hollow concrete blocks; and c, d vertical hollow clay brick units

Fig. 4.3 Historical evolution of the masonry units used in the building’s façade in Portugal

are capable of simulating the IP and OOP behaviour interaction [3–6]. The major advantage of the strut modelling approaches is their simplicity to be implemented in the analysis of an entire building structure. The capacity of simulating the global response of panels and the corresponding influence in the structural response is obviously an important advantage since it requires lower computational effort.

86

4 Experimental Characterization of the As-Built Masonry …

However, there is a lack of experimental data concerning the mechanical properties of the infill panels to calibrate the numerical models with good accuracy. Additional experimental data about the material and mechanical properties of infill walls made with different masonry units or mortar are needed to provide useful information for the structural designer and to calibrate numerical models. Some experimental studies are available in the literature on experimental study of infill panels’ mechanical properties [7–10], however it is still reduced the amount of information due to the large variability of materials used in the construction, particularly in what concerns the infill panels made with HCHB units. Based on this lack of motivation, an extensive experimental campaign was carried out in the framework of the present thesis to achieve the material, mechanical and dynamic properties of infill masonry walls built with HCHB units. This experimental campaign was divided into three stages, namely: Stage 1: Material characterization of masonry units and mortar that were used to build the infill panels in the laboratory. According to the standard NP-EN 771–1 [11], compression strength tests were made in HCHB units with two different thicknesses and from two different suppliers. A total amount of 40 tests were carried out. Details regarding the test setup, loading protocol, failure mode and stress results will be presented and discussed in Sect. 5.2.2. Concerning the mortar tests, compression strength and flexural strength tests were carried out according to the standard NP EN 1015–11 [12]. A specific testing campaign to characterize the properties of mortar used to build the panels tested within this work and the respective results will be shown in the section corresponding to each type of infill panels’ tests for a better correlation with the obtained test results. The testing methodology will be presented in Sect. 5.2.3. Stage 2: Mechanical characterization of as-built wallet specimens. Four different types of experimental tests were carried, namely: compressive strength tests according to EN1052-1 [13], diagonal tensile strength tests according to ASTM E 519–02 [14] and RILEM TC 76-LUM [15] and flexural strength tests parallel and perpendicular to the horizontal bed joints according to EN 1052–2 [16]. Additionally, it were also tested panels that were collected from existing infill walls (also built in the laboratory). These wallets were subjected to the same mechanical characterization tests and their results were compared with the ones obtained by the as-built specimens. A total amount of 64 tests were carried out. Details regarding the specimens’ geometric dimensions, material properties, test setup and instrumentation are described for each type of test. The results are presented and discussed in terms of stress–strain curves and damages observed within the tests. Finally, the following parameters will be also assessed, namely: (a) differences between the mechanical properties of the infill panels built with HCHB units with different thicknesses, (b) influence of plaster, and (c) differences between the mechanical properties of the existing and the as-built specimens.

4.1 Framework

87

As-built infill walls

Stage 2

Stage 1

Masonry unit

Compression strength tests NP EN 771-1

Mortar

Compression strength tests NP EN 1015-11

Flexural Compressive Diagonal tensile strength tests strength tests strength tests NP EN 1052-11 ASTM E 519-02 parallel to the RILEM TC 76 horizontal bed joints NP EN 1015-11

Stage 3

Flexural strength tests perpendicular to the horizontal bed joints NP EN 1015-11

Laboratory

In-situ

Flexural strength tests NP EN 1015-11

Fig. 4.4 Schematic layout of the testing campaign for material, mechanical and dynamic characterization

Stage 3: Ambient vibration tests carried out in laboratory and in-situ to characterize the infill panels’ dynamic properties, namely the IP and OOP vibration modes and corresponding natural frequencies. First, ambient vibration tests were performed in laboratory where the methodology to adopt in-situ was developed. These tests were carried out in a full-scale infilled frame, which allowed studying the IP and OOP frequencies and vibration modes. Additionally, the evolution of natural frequencies during the period after construction and the impact of different gravity load levels (applied in the adjacent columns) in the panel OOP frequencies were also analysed. Finally, ambient vibration tests were carried out in infill panels built with HCHB units with different thicknesses, which should allow to understand of the effect of the slenderness in the panels’ frequencies. The second part of this Stage 2 comprises a set of ambient vibration tests performed in-situ, in three different RC buildings, where the influence of the following parameters in the infill panels’ OOP frequencies was studied: geometric dimensions; boundary conditions and openings. Concerning this stage, it will be presented and discussed the major test results and findings as well as the details regarding each infill wall studied, the testing methodology and setups. Figure 4.4 presents the schematic layout of the overall testing campaign aiming to characterize the material, mechanical and dynamic properties of the infill walls components.

88

4 Experimental Characterization of the As-Built Masonry …

4.2 Stage 1—Material Characterization Tests of Masonry Units and Mortar 4.2.1 Introduction Stage 1 is related to the testing and characterization of the masonry units and mortar. Hollow Clay Horizontal Brick (HCHB) units 110 mm and 150 mm thick, here designated HCHB110 and HCHB150, respectively, were selected for this study since they are the most common masonry units used in the Portuguese building stock (Fig. 4.5). Concerning the material properties of each type of masonry unit, Table 4.1 provides the main characteristics that can be found in the datasheet provided by the manufacturer (Preceram), which supplied the masonry units for the experimental works developed within this Thesis. From the Table 4.1, it can be found that the compressions strength of both masonry units are equal. The thermal transmission coefficient (U) of the HCHB150 is 1.45 times higher than that of HCHB110 unit. Lower differences are noticed regarding

200

200

b)

a)

Fig. 4.5 Geometric dimensions of the masonry units used in this experimental campaign: a HCHB110 and b HCHB150

Table 4.1 Summary of the masonry units’ material properties (information provided by the product datasheet) Material properties

HCHB110

HCHB150

Compressive strength fc,unit (MPa)

≥1.5

≥1.5

Content of active soluble salts

S0 category

S0 category

Fire reaction

A1 Euroclass

A1 Euroclass

Masonry unit mass (Kg/unit)

3.9

5.2

Thermal transmission coefficient U (m2 K/W)

0,29

0,42

Acoustic insulation Rw (dB)

40

43

% of voids

57.7

57.9

4.2 Stage 1—Material Characterization Tests of Masonry Units and Mortar

89

acoustic insulation. Finally, the percentage of voids (% of voids) is very similar between both units. In this stage, eight HCHB110 units and ten HCHB150 units, supplied by PRECERAM (herein designated supplier A) were tested. In addition, ten HCHB110 units and ten HCHB150 units supplied by a different Portuguese manufacturer (herein designated supplier B) were also tested. The main aim was to perform compression strength tests according to the standard NP EN 771-1 [11] and to compare the results among the masonry units in order to assess the variability of the results. The mortar used to build the infill panels within the present work is a ready mix of traditional mortar type M5 (“Ciarga”) which was considered a suitable choice concerning the normal practice of the Portuguese construction industry in the 1970s.

4.2.2 Compression Strength Tests of Masonry Units 4.2.2.1

General Considerations and Setup

The main scope of this type of test is to obtain the compressive strength of masonry units and basically consists in the application of a distributed vertical compressive loading in the top surface of the masonry units until reaching the failure. Throughout the test, the compressive force (with constant increment) is recorded and the loading application area is also measured for each specimen, following the procedures detailed in NP EN 771-1 [11] standard. For these tests, it was used a servo-hydraulic actuator with a maximum capacity of 100 kN (±100 mm) (Fig. 4.6). The remaining test setup used a stiff steel plate attached to the actuator to distribute the vertical loading throughout the specimen top surface. A hinged connection between the hydraulic actuator and the steel plate was adopted to accommodate possible irregularities or eccentricities. The verticality of the specimen and the load application was ensured for each specimen. During the testing, besides the measurement recorded by the internal Linear Variable Differential Transformer (LVDT) of the servo-hydraulic actuator, no additional instrumentation was used to monitor the vertical displacements of the specimens. Due to that, the elasticity modulus was not determined for these tests.

4.2.2.2

Specimens Dimensions

Additionally, the complete report each specimen dimensions was collected to allow calculating, with precision, the bricks compressive strength. The global dimensions of the bricks were determined as well as the thicknesses of the interior septs for both types of brick units, as shown in Fig. 4.7. The geometric dimensions of HCHB150 from group A, HCHB150 from group B, HCHB110 from group A and HCHB110 from Group B are shown in Tables 4.2, 4.3, 4.4, 4.5, 4.6 and 4.7, respectively. In each table the average, standard deviation (SD) and coefficient of variation (C.o.V.) were

90

4 Experimental Characterization of the As-Built Masonry …

a)

b)

Fig. 4.6 Compression strength tests of masonry units: test setup a HCHB110 units; and b HCHB150 units

a)

b)

Fig. 4.7 Measurement of the brick units’ geometric dimensions: a HCHB150 units; and b HCHB110 units

determined for each dimension (Lu , hu , tu , VL, VC, VD, VR, H0, H1, H2, H3 and H4). The dimensions VL, VC, VD, VR were obtained by analysing the respective group ViL, ViC, ViD, ViR.

4.2 Stage 1—Material Characterization Tests of Masonry Units and Mortar

91

Table 4.2 HCHB150-Group A: geometric dimensions Average (mm)

Lu

hu

tu

299.5

187.8

149.6

VL

VC

8.2

VD

6.6

VR

6.8

7.6

H0

H1

H2

9.3

6.7

6.6

H3 8.1

H4 9.4

SD (mm)

1.2

1.1

0.7

0.8

0.9

0.5

0.9

0.4

0.3

0.5

1.5

0.5

C.o.V. (%)

0.4

0.5

0.4

10.2

14.6

7.8

12.1

4.5

3.8

7.1

18.8

5.8

H3

Table 4.3 HCHB150-Group B: geometric dimensions (units in millimetres) Average (mm)

Lu

hu

tu

VL

289.9

188.9

146.5

7.1

VC

VD

6.2

VR

6.6

7.2

H0

H1

H2

7.9

6.5

6.4

7.3

H4 8.1

SD (mm)

0.6

1.3

2.6

0.4

0.7

0.8

0.7

0.4

0.3

0.3

0.9

0.3

C.o.V. (%)

0.2

0.7

1.8

6.1

12.3

12.2

10.4

4.4

4.8

5.3

13.4

4.1

H3

H4

Table 4.4 HCHB110-Group A: geometric dimensions (units in millimetres) Lu

hu

tu

VL

VC

VR

H0

H1

H2

286.4

186.6

106.6

7.8

6.3

7.7

7.4

6.1

5.7

5.9

7.8

SD (mm)

2.5

1.8

1.3

0.7

0.5

0.7

0.4

0.6

0.7

0.5

0.4

C.o.V. (%)

0.9

1.0

1.2

8.9

7.6

8.7

5.4

10.7

11.3

8.3

5.2

Average (mm)

Table 4.5 HCHB110-Group B: geometric dimensions (units in millimetres) Lu Average (mm)

hu

tu

VL

VC

VR

H0

H1

H2

H3

H4 10.1

295.2

185.7

106.7

10.3

7.8

8.7

10.4

6.9

6.6

6.5

SD (mm)

0.4

0.6

1.3

1.2

0.9

0.6

0.8

0.5

0.6

0.3

0.6

C.o.V. (%)

0.1

0.3

1.3

11.9

11.8

7.4

8.1

7.1

8.4

4.7

5.5

4.2.2.3

Experimental Results

From the compression strength tests, the damages observed were characterized by progressive cracking and spalling parts of the bricks, which were revealed to be fragile as expected. For low loading demands, longitudinal cracks in all the horizontal septs are visible. Figure 4.8 shows pictures of the test stages for masonry units HCHB150. The brick side where cracking started developing, shown in Fig. 4.8a, conditioned by geometric defects of the brick during production. After that, the spalling of half-brick occurred, and consequently at the force drops the same time (Fig. 4.8b). Afterwards, the force re-started to increase until the collapse of another vertical sept (Fig. 4.8c). Finally, the last vertical sept falls for low force demands, which is expectable due to the high instability caused by the absence of horizontal vertical septs (Fig. 4.8d). The same behaviour was observed in the brick units from both groups.

92

4 Experimental Characterization of the As-Built Masonry …

Table 4.6 Summary of the experimental results: HCHB150 (Group A) Specimen

Gross area (mm2 )

Fi (N)

δ 1.1

fb,i (MPa)

Effective area (mm2 )

% voids

150_1_A

45,000

47,422

1.16

9180

79.6

150_2_A

45,000

58,001

1.42

9293

79.4

150_3_A

44,700

26,592

0.65

9171

79.5

150_4_A

44,700

54,992

1.35

8903

80.1

150_5_A

44,550

43,273

1.07

8502

80.9

150_6_A

44,252

46,701

1.16

8581

80.6

150_7_A

45,000

28,885

0.71

8100

82

15_8_A

45,000

45,103

1.10

9015

79.9

15_9_A

45,150

33,187

0.81

8533

81.1

15_10_A

44,700

37,830

0.93

8280

81.5

Average

1.04

Average

80.5

S.D.

0.24

S.D.

0.86

C.o.V. (%)

1.06

C.o.V. (%)

23.6

Tables 4.6 and 4.7 present the summary of the experimental results obtained in the compression strength tests in the brick units HCHB150 of Groups A and B, respectively. The gross area Ai was calculated based on the surface area where the compressive force was applied. The force, Fi , is the peak load reached in each test and the δ coefficient was determined according to Table A1 of ANNEX A of standard NP EN 772-1 [11]. The compression strength of each masonry unit, fb,i , was then computed according to the code standard as expressed by Eq. 4.14. Finally, the effective area consisting of the vertical septs was computed as well the percentage of voids of each brick unit.

fb,i =

Fi ×δ Ai

(4.14)

From the results, it can be pointed out that Group A achieved a mean compression strength equal to 1.04 MPa, 1.24 times higher than Group B. The percentage of voids are quite similar between both groups, namely 80.5% and 81.5% concerning group A and B brick units, respectively. The supplier indicated that the minimum compression strength was 1.5 MPa, as presented in Table 4.1. However, the value is 1.44 times higher than the result herein obtained. Finally, the standard deviation (SD) and coefficient of variation (C.o.V.) was computed and similar values were found. The testing of the HCHB110 units revealed a similar behaviour (Fig. 4.9), namely characterized by the fragile behaviour and development of vertical cracks along the longitudinal alignments (Fig. 4.9a) and respective spalling (Fig. 4.9b), progressively with the increase of vertical force. Finally, after half-brick spalling, there is only one

4.2 Stage 1—Material Characterization Tests of Masonry Units and Mortar

a)

b)

c)

d)

93

Fig. 4.8 Compression strength tests of masonry units: HCHB150 damage evolution a cracking in the horizontal septs; b spalling of half-part of the brick; c collapse of three vertical septs; and d total rupture

vertical sept resisting to the vertical force until cracking and corresponding rupture (Fig. 4.9c, d). The behaviour of the specimens of both Groups A and B was similar. The summary of the compression strength tests in HCHB110 brick units are shown in Tables 4.8 and 4.9. The compression strength results achieved an average value equal to 1.08 MPa and 1.73 MPa respectively, for Groups A and B. For these units, higher differences can be observed, since the Group B units reached a value

94

4 Experimental Characterization of the As-Built Masonry …

Table 4.7 Summary of the experimental results: HCHB150 (Group B) Specimen

Gross area (mm2 )

F (N)

δ

150_1_B

42,050

28,031

1.1

150_2_B

42,340

37,337

150_3_B

42,486

150_4_B 150_5_B

fb,i (MPa)

Effective area (mm2 )

% voids

0.73

7917

81.2

0.97

8149

80.7

23,874

0.62

8083

80.9

43,210

33,511

0.85

7895

81.7

42,194

32,690

0.85

7724

81.7

150_6_B

42,050

36,313

0.95

7743

81.6

150_7_B

43,500

26,974

0.68

7547

82.7

150_8_B

40,890

46,825

1.26

7961

80.5

150_9_B

43,061

40,513

1.03

7962

81.5

150_10_B

42,920

24,824

0.64

7395

82.8

Average

0.86

Average

81.5

S.D.

0.19

S.D.

0.70

C.o.V. (%)

0.85

C.o.V. (%)

22.3

1.60 times higher, however with higher variation (29.4% compared with the 19.7%). These results can be justified by the lower percentage of voids of the Group B (larger thickness of the vertical septs), 10% higher, and consequent larger area to resist to the vertical force. By comparing the HCHB110 and HCHB150 units, it can be found that the compression strength of the thicker bricks is higher than the larger ones, which can be justified by: (i) larger thickness of the vertical septs; (ii) lower defects such as micro-cracking due to the construction process. Table 4.10 shows the summary of the results herein obtained, where it can be easily compared the results from each group and better assess the differences among them.

4.2.3 Compression and Flexural Strength Tests of the Mortar The infill wallets tested in this research thesis were built using an industrial readymix mortar of M5 class (“Ciarga” type). Samples of the mortar used in the specimens ‘construction were collected in order to assess their compressive and tensile strength according to the standard EN 1015-11 [12], as shown in Fig. 4.10. Specimens were collected for each group of tests carried out in this thesis to understand better the possible variables that could affect the results obtained. Thus, in each group of tests, such as mechanical characterization tests and full-scale OOP tests, the mortar properties obtained from the compressive strength and tensile strength according to EN 1015-11 [12] will be shown in the sub-section of each type of test.

4.2 Stage 1—Material Characterization Tests of Masonry Units and Mortar

a)

b)

c)

d)

95

Fig. 4.9 Compression strength tests of masonry units: HCHB110 damage evolution a cracking in the horizontal septs; b spalling of half-part of the brick; c collapse of two vertical septs; and d total rupture

The mortar properties used to build the wallets of the mechanical characterization tests are summarized in Sect. 5.3.1. The mortar properties used to build the full-scale infill panels for the OOP tests are presented Chap. 6.

96

4 Experimental Characterization of the As-Built Masonry …

Table 4.8 Summary of the experimental results: HCHB110 (Group A) Specimen

Gross area (mm2 )

F (N)

δ 1.1

fb,i (MPa)

Effective area (mm2 )

% voids

110_1_A

31,320

33,211

1.17

6648

78.8

110_2_A

29,925

25,796

0.95

6327

78.9

110_3_A

30,888

20,936

0.75

6421

79.2

110_4_A

30,210

25,579

0.93

6199

79.5

110_5_A

30,709

26,266

0.94

5869

80.9

110_6_A

31,320

37,231

1.31

6315

79.8

110_7_A

29,715

38,734

1.43

6212

79.1

110_8_A

30,210

32,184

1.17

5935

80.3

1.08

Average

81.5

0.21

S.D.

0.70

C.o.V. (%)

0.85

Average S.D. C.o.V. (%)

19.7

Table 4.9 Summary of the experimental results: HCHB110 (Group B) Specimen

Gross area (mm2 )

F (N)

δ

110_1_B

31,672

59,453

1.1

110_2_B

31,270

68,376

110_3_B

30,975

110_4_B

31,968

110_5_B

Effective area (mm2 )

% voids

2.06

8096

74.4

2.41

7183

77.1

48,649

1.73

8459

72.7

35,699

1.23

8591

73.1

30,975

62,821

2.23

8135

73.7

110_6_B

31,860

62,821

2.17

6771

78.7

110_7_B

31,860

18,136

0.63

7899

75.2

110_8_B

31,565

41,597

1.45

8009

74.6

110_9_B

30,975

46,291

1.64

8201

73.5

110_10_B

31,860

49,723

1.72

7596

76.2

1.73

Average

74.9

0.51

S.D.

1.80

C.o.V. (%)

2.40

Average S.D. C.o.V. (%)

fb,i (MPa)

29.4

Table 4.10 Summary of the results obtained in the compression strength tests of brick units Brick unit

fb,ave (MPa)

SD (MPa)

C.o.V. (%)

% of voids (average)

HCHB150 (Group A)

1.04

0.24

23.6

80.5

HCHB150 (Group B)

0.86

0.19

22.3

81.5

HCHB110 (Group A)

1.08

0.21

19.7

81.5

HCHB110 (Group B)

1.73

0.51

29.4

74.9

4.3 Stage 2—Mechanical Characterization of the Infill Masonry Wallets

a)

97

b)

Fig. 4.10 Mortar samples for testing according to EN 1015–11 [12]: a compression strength test; and b flexural strength test

4.3 Stage 2—Mechanical Characterization of the Infill Masonry Wallets 4.3.1 Testing Campaign Overview The testing campaign here described aims to provide complete insight regarding the mechanical properties of infill panels built with HCHB brick units, namely in terms of their compressive strength, elastic modulus, diagonal tensile strength, shear modulus and flexural strength perpendicular and parallel to the horizontal joints. Specimens consisting of two different types of infill masonry wallets comprised the experimental campaign, namely: Infill wallet type 1: Panel built with HCHB units. The selected clay brick units for this specific study are the horizontal hollow clay bricks 110 mm and 150 mm thick, supplied by PRECERAM (Group A of Sect. 4.3.2). Infill wallet type 2: panels collected from existing infill walls in the lab (herein designated as “existing” wallets) and made with HCHB150 units, with or without plaster. Vertical and horizontal cuts were performed by technical staff specialized in demolitions. After normalized the specimen dimensions, a top layer of mortar was included to regularize the top and bottom parts of the specimens (Fig. 4.11a, b). Concerning the material properties of each masonry unit type, the major properties were presented, detailed and discussed in the previous Sect. 5.2. No plaster was used for the specimens built in the laboratory with HCHB110 brick units. Eleven specimens made with HCHB150 units were built with 10 mm plaster thickness (only in one side of the specimen). The specimens with plaster, here designated as HCHB150P10, were only tested to evaluate the flexural strength. The aim was to compare the flexural strength of both wallets, with and without plaster. The construction methodology adopted to build the specimens with HCHB units was following the current practice in Portugal. For all these specimens, a vertical and horizontal bed joints were made with approximately 10 mm thickness (Fig. 4.12).

98

4 Experimental Characterization of the As-Built Masonry …

a)

b)

Fig. 4.11 Details of the existing masonry panels made with HCHB150 units: a front view; and b lateral views

Fig. 4.12 Example of construction of wallets made with HCHB110 units

4.3.2 Mechanical Properties of Mortar The experimental results from mortar testing are summarized in Table 4.11 according to each specimens’ group and the respective type of experimental test that they were subjected to. The compressive strength (fc,mo ) and tensile strength (ft,mo ) are shown

4.3 Stage 2—Mechanical Characterization of the Infill Masonry Wallets

99

Table 4.11 Summary of the mortar material properties of each specimens’ group

Type of wallets

Group of wallets

I

II

III

IV

Mortar property

fc,mo (MPa)

ft,mo (MPa)

fc,mo (MPa)

ft,mo (MPa)

fc,mo (MPa)

ft,mo (MPa)

fc,mo (MPa)

ft,mo (MPa)

HCHB110

8.08

2.75

8.08

2.75

10.61

3.28

8.08

2.75

HCHB150

13.35

4.51

12.38

4.76

14.13

4.98

14.13

4.98

HCHB150P10

N/A

N/A

N/A

N/A

12.95

4.52

12.95

4.52

HCHB150 (existing)

11.90

9.03

11.90

9.03

11.90

9.03

11.90

9.03

HCHB150P10 (existing)

12.60

5.18

N/A

N/A

12.60

5.18

N/A

N/A

N/A—Not applicable

for each type group of masonry wallets. From the results, slight variations can be found, which can be justified by the workmanship and by the experimental variability. Mortar samples were grouped according to each group of wallets in which they were used and the corresponding type of test, namely: Group I—Compressive strength tests Group II—Diagonal tensile strength tests Group III—Flexural strength tests parallel to the horizontal bed joints Group IV—Flexural strength tests perpendicular to the horizontal bed joints A total amount of 80 specimens were tested in this experimental campaign. Four different types of tests were carried out, namely compressive strength tests using 3 specimens, the minimum number required by the standard EN1052-1 [13], diagonal tensile strength tests also in 3 specimens, (the minimum according to ASTM E 519-02 [14], flexural strength tests parallel and perpendicular to the horizontal bed joints (for which the number minimum of specimens according to EN 1052-2 [16] is 5 of each type). The number of tested specimens for each set of tests is summarized in Table 4.12.

4.3.3 Compressive Strength Tests 4.3.3.1

General Considerations

The compressive strength tests were carried out to achieve the compressive strength and the elastic modulus perpendicular to the horizontal bed joints according to the standard EN 1052-1 [13]. The specimens’ geometric dimensions were defined according to the standard EN 1052-1 [13] recommendations, summarized in Table 4.13 and shown in Fig. 4.13. Where lu and hu are the masonry unit length and height, respectively; ls , hs and ts are the masonry wallets length, height and thickness.

100

4 Experimental Characterization of the As-Built Masonry …

Table 4.12 Summary of the experimental campaign: number of specimens tested Type of Compressive masonry wallets strength tests

Diagonal tensile strength tests

Flexural strength tests parallel to the horizontal bed joints

Flexural strength tests perpendicular to the horizontal bed joints

HCHB110

5

4

5

5

HCHB150

5

5

5

5

HCHB150P10

N/A

N/A

6

5

HCHB150 (existing)

4

1

1

3

HCHB150P10 (existing)

4

N/A

1

N/A

Total

18

10

18

18

HCHB110—wallets made with HCHB with 110 mm thickness and no plaster HCHB150—wallets made with HCHB with 150 mm thickness and no plaster HCHB150P10—wallets made with HCHB with 150 mm thickness with 10 mm plaster HCHB150 (existing)—wallets made with HCHB with 150 mm thickness and no plaster taken from existing full-scale wall panels in Lab HCHB150P10 (existing)—wallets made with HCHB with 150 mm thickness, with 10 mm plaster, taken from existing full-scale wall panels in Lab

Table 4.13 Masonry wallets’ geometric dimensions according to the standard EN 1052–1 [13] recommendations Masonry unit dimensions

Masonry wallet geometric dimensions

lu (mm)

hu (mm)

Length ls (mm)

Height hs (mm)

≤300

≤150

≥(2 × lu )

≥5 hu

>300

≤150

≥(1.5 × lu )

≥5 hu

≥3 hu

>150 >150

Thickness ts (mm) ≥3 ts and ≤15ts

≥tu

≥3 hu

This test consists of applying a distributed vertical compressive load in the top surface of the specimen until the failure is reached. For this, a servo-hydraulic actuator with a maximum capacity of 100 kN (±100 mm) was used, as shown in Fig. 4.14. As for the test setup, a steel shape was attached to the actuator to distribute the vertical loading throughout the specimen top surface. A hinged connection between the hydraulic actuator and the steel shape was assumed to accommodate possible irregularities or eccentricities. The specimens’ dimensions were 600 × 600 mm, wide and height, respectively (Fig. 4.15). The instrumentation was assumed according to the standard EN 1052-1 [13], consisting of two pairs of vertical LVDT’s and one horizontal transducer in each specimen face. The layout of the instrumentation used for each type of specimen is shown in Fig. 4.15.

4.3 Stage 2—Mechanical Characterization of the Infill Masonry Wallets

101

Fig. 4.13 Compressive strength tests: specimens’ geometric dimensions according to standard EN 1052–1 [13]

4.3.3.2

Experimental Results

Stress-Strain Curves The compressive strength of each specimen was evaluated according to EN 1052-1 [13], as the maximum force applied until failure occurs. The force was divided by

102

4 Experimental Characterization of the As-Built Masonry …

a)

b)

Fig. 4.14 Compressive strength tests: test setup a Front view; and b Lateral view

(a)

(b)

Fig. 4.15 Compressive strength tests: specimens’ geometric dimensions and respective instrumentation

4.3 Stage 2—Mechanical Characterization of the Infill Masonry Wallets

103

the area in which was applied the compression loading (in the case of the specimens with plaster HCHB150P10 and HCHB150P10 (existing) the area considered for the calculation was the surface area of the brick thickness plus the plaster thickness). From a set of experiments, the characteristic compressive strength (fc,k ) can be taken as the lower value of two conditions: (1) fc,mean /1.2 (where fc,mean stands for the mean compressive strength of a set of experiments; and (2) fi,min which is the minimum compressive strength obtained by a specimen in a set of experiments. For each specimen it was also calculated the elastic modulus according to the standard EN 1052-1 [13] suggestion, which consists of the computation of the secant elastic modulus for a stress level equal to 1/3 of the maximum stress and the corresponding average strain was taken from all vertical LVDTs measurements. The stress-strain curves obtained for each group of specimens are illustrated in Fig. 4.16. The mean, standard deviation and coefficient of variation values were also computed for each group. Additionally, the average curve (μ) of the stress-strain results obtained by all the specimens, the average curve plus two times the standard deviation (μ + 2SD) and the average curve minus two times the standard deviation (μ−2SD) were included for each Group result. From the strain-stress curves, the following observations can be drawn: • The lowest compressive strength was reached by the HCHB110 group with 0.66 MPa, as shown in Fig. 4.16a; • The HCHB150 group achieved the highest mean compressive strength equal to 1.09 MPa and a lower C.o.V. of 12.81% (Fig. 4.16b); • By comparing the results among the HCHB110 and the HCHB150 groups, it can be observed that the last one is 40% higher than the result obtained by the HCHB110 group. On the other hand, larger variability was found in the HCHB110 specimens’ tests. From the comparison between the HCHB150 and HCHB150 (existing) groups, it can be observed that the last group reached a mean compressive strength equal to 0.81 MPa, 18% lower than the HCHB150 but with similar coefficient of variation. Besides the material variability and experimental variability no specific reason can justify this difference. Moreover, the mortar properties of the HCHB150 specimens is slightly higher than that used in HCHB150 (existing) specimens, shown in Fig. 4.16c. Some minor cracking caused in the preparation of the HCHB150 (existing) specimens may affect the results. The HCHB150P10 (existing) group reached a mean compressive strength of 0.89 MPa (Fig. 4.16d), which is 10% higher when compared with the results of the HCHB150 (existing) group results. Additionally, Fig. 4.16e includes the plot of the average curves obtained for each group of tests in which some differences among them can be detected. For example, for a vertical strain of 0.25% the HCHB150P10 (existing) group reached a mean compressive strength of 0.44 MPa, 1.62 times higher than the HCHB150 (existing) group, 1.91 times higher than the HCHB150 and 2.2 times higher than the HCHB110 group. For a vertical strain equal to 0.45%, the HCHB150P10 (existent)

104

4 Experimental Characterization of the As-Built Masonry …

(a)

(b)

(c)

(d)

(e) Fig. 4.16 Compressive strength test results: Stress versus Strain curves a HCHB110, b HCHB150, c HCHB150 (existing); d HCHB 150P10 (existing); and e global average curves

group again obtained the maximum compressive strength, 1.4 times higher than the group HCHB150 (existent), 1.8 times higher than the group HCHB150 and 2.18 times higher than HCHB110 one. As future work, the present experimental campaign must be completed with compression strength tests to be carried out in specimens made with HCHB 150 and HCHB 110 units with 10 mm plaster, thus performing the groups HCHB150P10 and HCHB110P10, respectively. This will allow an understanding of the plaster’s contribution to the infill panels’ compressive strength capacity. Table

4.3 Stage 2—Mechanical Characterization of the Infill Masonry Wallets

105

Table 4.14 Compressive strength results: statistical parameters Statistical parameter

HCHB110

HCHB150

HCHB150 (existing)

HCHB150P10 (existing)

fc,mean (MPa)

0.66

1.09

0.81

0.89

SD (MPa)

0.13

0.14

0.10

0.18

C.o.V (%)

19.68

12.81

12.62

19.68

fc,k (MPa)

0.54

0.91

N/A

N/A

N/A—Not applicable

Table 4.15 Elasticity modulus results: statistical parameters Statistical parameter

HCHB110

HCHB150

HCHB150 (existing)

HCHB150P10 (existing)

Emean (MPa)

1837

1975

2067

2221

SD (MPa)

563

719

286

324

C.o.V. (%)

30.60

36.70

15.20

14.50

Emean /fc,mean

2783

1811

2564

2484

4.14 summarises the compressive strength statistical parameters obtained by each group of specimens. The statistical parameters obtained for the specimens’ elastic modulus are summarized in Table 4.15. From those results, it was found that the highest result was achieved by the HCHB150P10 (existing) group with 2221 MPa and the lowest one by the HCHB110 group with 1837 MPa. Slight differences can be observed between the HCHB150 (existing) group and the HCHB150 (about 5% higher), which can be justified by the presence of mortar inside of some of the brick holes of the HCHB150 (existing) specimens that may have contributed to the increase of the wallets stiffness. Since those specimens were extracted from existing wall panels, some pieces of mortar from the vertical joints that filled some holes could not be removed. This detail may have provided additional stiffness to the specimens. By comparing the results of the existing specimens with and without plaster, it can be observed that the plaster contributed to an increase of the elastic modulus of around 7%. Lastly, the HCHB150 reached a mean elastic modulus about 8% higher than HCHB110 group. Additionally, it was computed the ratio between the infill panels’ elastic modulus and the compressive strength (Emean /fc,mean ). The results are summarized again in Table 4.15, from which it can be found that the highest value was achieved by the HCHB110 group and the lowest by the HCHB150 group with a ratio equal to 2783 and 1811, respectively. The results obtained by the HCHB150 group are 35% lower than those from the group HCHB110. From the comparison between the ratios obtained by the groups HCHB150 and HCHB150 (existing), it can be observed that the existing one is 30% higher, which can be related to the combination of lower compressive strength and higher elastic modulus. Concerning only the existing specimens, the plaster did not contribute to higher Emean /fc,mean ratio.

106

4 Experimental Characterization of the As-Built Masonry …

Looking for some results available in the literature, it can be found that [17] carried out compressive strength tests in small infill panels made with HCHB150 units, which obtained a mean compressive strength equal to 1.26 MPa (13% higher), an elastic modulus equal to 1577 MPa (13% lower) and a ratio Emean /fc,mean equal to 1252. This result is 31% lower than the one obtained by the HCHB150 group within this testing campaign. Through the literature, it is possible to find some analytical relationships between the infill panels’ compressive strength and their elastic modulus that different authors proposed. In 1971, [18] proposed the relationship expressed by the Eq. 4.15: Em = 750fc

(4.15)

Em = 1180fc0.83

(4.16)

While [19] proposed Eq. 4.16:

Later, [20] proposed the Eq. 4.17: Em = 500fc

(4.17)

Em = 2116fc0.50

(4.18)

Hendry [21] reported Eq. 4.18:

Due to the large dispersion of results obtained from previous analytical predictions when compared with the experimental results, Paulay et al. [22] and [18] converged in the same expression (Eq. 4.19), which is also suggested by Eurocode 6 [23] in Sect. 3.8.3: Em = 1000fc

(4.19)

Large differences can be found by comparing the prediction results according to the Eurocode 6 [23] proposal. The results of the groups HCHB110 and HCHB150 are 2.8 and 1.7 times higher than that obtained by the analytical prediction. This result allows concluding that this expression should probably be revised and adapted for infill panels made with different types of masonry units. Additional tests should be performed to create a database with enough data to extract analytical correlations between the compressive strength and the elastic modulus.

Observed Damages From the observation of the specimens’ damages throughout the tests, it was observed a fragile brittle behaviour in all the specimens, due to the quick development of cracks followed by sudden ruptures and spalling of vertical septs, or even the partial

4.3 Stage 2—Mechanical Characterization of the Infill Masonry Wallets

a)

c)

107

b)

d)

Fig. 4.17 Compressive strength tests: observed damages a HCHB110, b HCHB150; c HCHB150 (existing) and d HCHB150P10 (existing)

collapse of bricks without prior visible cracks propagation (Fig. 4.17). The HCHB110 specimens’ failure was characterized by the crushing of the top or bottom row of bricks, which caused the panel instability. Similar failures were observed in the remaining HCHB (existing) groups.

108

4 Experimental Characterization of the As-Built Masonry …

4.3.4 Diagonal Tensile Strength Tests 4.3.4.1

General Considerations

Diagonal tensile strength tests were carried out according to the American standard ASTM E 519-02 [14] and RILEM TC 76 [15]. This test consisted of applying a diagonal compressive force through two steel loading shoes in the top, and bottom part of a specimen rotated 45º relative to the vertical direction. The top shoe is attached to a servo-hydraulic actuator with a 1500 kN capacity (±75 mm), allowing continuous loading until the specimen reaches the rupture. Both standard requirements are unanimous concerning the specimens’ geometric dimensions (independent of the masonry unit thickness), which were taken as 1200 × 1200 mm width and height, respectively (Fig. 4.18). The minimum number of specimens required is 3. A total amount of 8 transducers were used to monitor the tests. It was adopted the instrumentation recommended by ASTM E 519-02 [14], to measure the vertical shortening strain (εv ) that was captured by a pair of LVDTs (one in each specimen’s face) and of the horizontal strain (εh ) by three pairs of LVDT (also three in each specimen’s face), as shown in Fig. 4.18b. The measurement length between the transducers used to calculate the shear distortion (U) was 1000 mm, as recommended by the standard. Additionally, two pairs of horizontal transducers were placed on both specimen’s faces with 600 mm in length. A total of ten tests were performed in this set of tests, four for HCHB110 group, five for HCHB150 group and one for the HCHB150 (existing) group. F/A

1000

600

1697 1000

600

600

a)

b)

Fig. 4.18 Diagonal tensile strength tests: a Test setup view and b instrumentation and specimen’s geometric dimensions

4.3 Stage 2—Mechanical Characterization of the Infill Masonry Wallets HCHB110 Group

Specimen 1 Specimen 2 Specimen 3 Specimen 4 μ curve μ−2SD curve μ+2SD curve

Shear Stress Ss (MPa)

0.7 0.6 0.5

0.7

0.4 0.3 0.2

Shear Stress: Ss,mean=0.57MPa

0.1

SD=0.20MPa C.o.V.=35.20%

0.0 -10

-8

-6

-4

-2

0

Vertical strain (%)

2

4

6

8

HCHB150 Group

0.8

Shear Stress Ss (MPa)

0.8

0.6 0.5

0.3

Shear Stress: Ss,mean=0.65MPa

0.2

SD=0.14MPa C.o.V.=22.20%

0.1

10 x10

0.0 -10

-4

Specimen 1 Specimen 2 Specimen 3 Specimen 4 Specimen 5 μ curve μ−2SD curve μ+2SD curve

0.4

-8

-6

-4

-2

Specimen 1

4

6

10 x10

8

-4

HCHB110 HCHB150 HCHB150 (existing)

0.8

Ss=0.38MPa 0.7

0.7

0.6

0.6

Shear Stress Ss (MPa)

Shear Stress Ss (MPa)

2

Horitontal strain (%)

b)

HCHB150 (existing) Group

0.5 0.4 0.3 0.2

0.5 0.4 0.3 0.2 0.1

0.1 0.0 -10

0

Vertical strain (%)

Horitontal strain (%)

a) 0.8

109

0.0

-8

-6

-4

-2

0

Vertical strain (%)

2

4

6

8

10 x10

-4

-6

-5

-4

-3

-2

Vertical strain (%)

Horitontal strain (%)

-1

0

1

-4 2 x10

Horitontal strain (%)

d)

c)

Fig. 4.19 Diagonal tensile strength test results: a HCHB110; b HCHB150; c HCHB150 (existing); and d Global groups’ comparison

4.3.4.2

Experimental Results

Stress-Strain Curves The shear stress (Ss ) was computed for each specimen according to the Eq. 4.20, provided by the standard ASTM E 519-02 [14], which consists in the relationship between the maximum applied force (F) and the cross-section transverse net area (An ). For each specimen, the shear-stress curves versus the vertical shortening and the horizontal extension are presented in Fig. 4.19. Additionally, each specimen’s transverse or shear stiffness (modulus of rigidity), G, was evaluated based on the shear stress-strain curves, as suggested by the standard [14]. The shear stress results of all the groups are summarized in Table 4.16.

SS =

0.707 × F An

(4.20)

The shear distortion, U, was evaluated according to the Eq. 4.21, where ΔV is the vertical shortening, ΔH is the horizontal extension and g is the vertical gage length, as recommended by [14].

110 Table 4.16 Shear stress results: statistical parameters

4 Experimental Characterization of the As-Built Masonry … HCHB110

Statistical parameter Ss,mean (MPa)

0.57

HCHB150 0.65

HCHB150 (existing) 0.38

SD (MPa)

0.20

0.14

N/A

C.o.V. (%)

35.20

22.20

N/A

N/A—Not applicable

γ =

ΔV × ΔH g

(4.21)

The average curve (μ) of the stress-strain results obtained by all the specimens, the average curve plus two times the standard deviation (μ + 2SD) and the average curve minus two times the standard deviation (μ−2SD) were included for each group result, plotted in Fig. 4.19. From the shear stress versus vertical and horizontal strain curves (see Fig. 4.19), the following observations can be drawn: • From the comparison between the groups HCHB110 and HCHB150 it can be observed that the first group achieved a mean shear stress of 0.57 MPa, 13% lower than the second group (Fig. 4.19a, b, respectively); • The HCHB 150 group (Fig. 4.19b) achieved the highest shear stress with a mean value of 0.645 MPa, while the lowest one was reached by the HCHB150 (existing) group with a mean value equal to 0.38 MPa (Fig. 4.19c) which is about 42% lower; • The result obtained by the sole specimen HCHB150 (existing), Fig. 4.19c, reached a shear stress of about 0.39 MPa, which is 41% lower than the HCHB150 result. Again, this can be justified by experimental variability or due to some minor cracking that affected the specimen performance. No solid conclusions can be taken since only one test was performed. Concerning the coefficient of variation, the HCHB110 group gathered the largest results with a coefficient of variation equal to 35.2%. From the shear stress-distortion curves plotted in Fig. 4.20, it was found that, the mean shear stiffness equal to 1141 MPa was obtained by the group HCHB110 (Fig. 4.20a) that is 13% higher than the HCHB150 group, for which the lowest shear stiffness of 996 MPa was achieved (Fig. 4.20b). The highest shear stiffness was obtained by the HCHB150 (existing) group (Fig. 4.20c) with a mean value equal to 1198 MPa, which is 1.2 times higher than the HCHB150 result (Fig. 4.20b). The ratio between the elastic modulus and the shear stiffness for each group and is shown in Table 4.17. From this, it was found that the highest result was obtained for the HCHB110 group with a ratio equal to 0.621 and the lowest one obtained for HCHB150 with a ratio equal to 0.481, which is 23% lower. Concerning the variability of shear stiffness results, it was found that the largest C.o.V. was achieved for HCHB110 group with a value equal to 11.8% and the smallest one obtained by the group HCHB150 with a C.o.V. equal to 8.91%. Figure 4.20d compares the global average curves obtained by each group. Similarly, Table 4.17

4.3 Stage 2—Mechanical Characterization of the Infill Masonry Wallets HCHB110 Group

Specimen 1 Specimen 2 Specimen 3 Specimen 4 μ curve μ−2SD curve μ+2SD curve

0.6

HCHB150 Group 0.8

Shear Stress Ss (MPa)

Shear Stress Ss (MPa)

0.8

0.4

Shear stiffness Gmean=1141MPa

0.2

0.6

0.2

0.4

0.6

0.8

1.0

1.2

Shear stiffness Gmean=996MPa

0.2

SD=88.70MPa CoV=8.91% 1.4x10

0.0 0.0

-4

0.2

Shear distortion γ (mm/mm)

0.4

0.6

0.8

1.0

1.2

1.4x10

-4

Shear distortion γ (mm/mm)

a)

b) 0.5

HCHB150 (existing) Group 0.8

Specimen 1 Specimen 2 Specimen 3 Specimen 4 Specimen 5 μ curve μ−2SD curve μ+2SD curve

0.4

SD=135MPa CoV=11.80% 0.0 0.0

111

G= 1198MPa

Shear Stress Ss (MPa)

Shear Stress Ss (MPa)

0.4

0.6

0.4

0.2

0.0 0.0

0.2

0.4

0.6

0.8

1.0

1.2

1.4x10

-4

0.3

0.2

HCHB110 HCHB150 HCHB150 (existing)

0.1

0.0 0.0

0.1

Shear distortion γ (mm/mm)

0.2

0.3

0.4

0.5

0.6x10

-4

Shear distortion γ (mm/mm)

c)

d)

Fig. 4.20 Shear stress results: a HCHB110; b HCHB150; c HCHB150 (existing); and d Global groups’ comparison

Table 4.17 Shear stiffness (G) modulus’ results: statistical parameters

Statistical parameter

HCHB110

HCHB150

HCHB150 (existing)

Gmean (MPa)

1141

996

1198

SD (MPa)

135

88.7

N/A

C.o.V. (%)

11.8

8.91

N/A

Gmean /Emean

0.621

0.481

0.531

N/A—Not applicable

summarizes the statistical parameters for the transverse shear stiffness modulus of each group. Again, the average curve, μ, of the stress-strain results obtained by all the specimens, the average curve plus two times the standard deviation, μ + 2SD, and the average curve minus two times the standard deviation, μ-2SD, were included for each Group result.

Observed Damages The most representative failure modes of each group are presented in Fig. 4.21. The tests’ observation found a common failure mode in all the tested specimens from

112

4 Experimental Characterization of the As-Built Masonry …

a)

b)

c) Fig. 4.21 Diagonal tensile strength tests: Failure modes a HCHB110; b HCHB150 and c HCHB150 (existing)

HCHB150 and HCHB150 (existing), which was characterized by the separation between the brick-joint interfaces (Fig. 4.21a, c). No pure vertical cracking was observed in any specimen. Some minor cracking was observed in the top and bottom parts of the specimens and, as in the compression strength tests, brittle behaviour was observed. Finally, it should be mentioned that sliding failure occurred at the middle height of the panel in two specimens of the group HCHB110, shown in Fig. 4.21b.

4.3 Stage 2—Mechanical Characterization of the Infill Masonry Wallets

113

4.3.5 Flexural Strength Tests 4.3.5.1

Flexural Strength Tests Parallel to the Horizontal Bed Joints

The flexural strength tests parallel to the horizontal bed joints were carried out according to the standard EN1052-2 [16]. This test type aims to obtain the wallets flexural strength according to the principal direction where the internal forces (bending moments and shear) develop. The standard NP EN1052-2 [16] specify the following requirements for the definition of the specimens’ dimensions (Fig. 4.22), namely: (1) Specimens’ width (b): b ≥ 400 mm and b ≥ 1.5xlu , where lu is the masonry unit length; (2) At least 2 horizontal joints must exist within the l2 dimension. For this, a servo-hydraulic actuator with a maximum capacity of 100 kN (±100 mm) was used. The loading was distributed along two linear alignments distanced 300 mm apart. The loading was applied through two rollers placed between the specimen and the steel shapes that were attached to the actuator. The rollers are free to rotate, thus allowing the rotation of specimen. The specimen reacted against a steel structure composed also of two horizontal rollers inserted in other steel shapes, as shown in Fig. 4.23. The OOP restraints were positioned according to the standard, 50 mm distanced from the specimens’ top and bottom sections.

Fig. 4.22 Flexural strength tests parallel to horizontal bed joints: specimens’ geometric dimensions according to NP EN1052-2 standard [16]

114

4 Experimental Characterization of the As-Built Masonry …

a)

b)

Fig. 4.23 Flexural strength tests parallel to horizontal bed joints: test setup for a HCHB150 specimen; and b HCHB110 specimen

The specimens’ geometric dimensions were defined as 450 × 600 mm, as shown in Fig. 4.24. The standard also requires a minimum number of 5 specimens whose rupture occurs between their OOP loading horizontal alignments. The standard excludes all the tests whose failure is different from the mentioned before. The standard provides no recommendations for the instrumentation. The standard only requires the load monitoring of the tests. However, this work monitored four points distanced from the geometric centre (50 mm vertical and 75 mm horizontal), covering a contour of 100 × 150 mm (Fig. 4.24).

Experimental Results Stress-Strain Curves The flexural strength was computed according to the standard EN 1052-2 [16] expression, given by Eq. 4.22: fb,parallel =

3 × fi,max × (l1 − l2 ) 2 × b × tu2

(4.22)

where, fi,max is the maximum applied force, l1 is the distance between OOP restrains, l2 is the distance between the internal OOP loading application alignments, b is the specimen width and tu is the specimen thickness. From the testing campaign, three specimens HCHB150 did not reach the failure according to the standard and

4.3 Stage 2—Mechanical Characterization of the Infill Masonry Wallets

a)

115

b)

Fig. 4.24 Flexural strength tests parallel to horizontal bed joints: a detail view of the instrumentation and b instrumentation and specimens’ dimensions

therefore were excluded from the results. The force-displacement curves are plotted in Fig. 4.25. The lowest flexural strength was reached by HCHB110 (Fig. 4.25a) with a mean value equal to 0.12 MPa, which is 15% lower than the flexural strength obtained by the group HCHB150 (Fig. 4.25b) that was equal to 0.14 MPa. The highest flexural strength was reached by the group HCHB150P10 (Fig. 4.25c) with a mean value equal to 0.22 MPa, that is 1.57 times higher than HCHB150 group. The HCHB150P10 (existing), shown in Fig. 4.25d, reached a flexural strength equal to 0.04 MPa, which significantly lower than HCHB150P10 group. From the global results analysis (Fig. 4.25e), it becomes evident that the plaster contributed to increased flexural strength. Concerning to the plaster effect in the as-built specimens, it was found that the specimens of the group HCHB150P10 reached a mean flexural strength of 0.218 MPa, meaning that the plaster increased the flexural strength by 57% when compared with the result obtained by the HCHB150 specimens; note that only 2 tests satisfied the criteria defined by the standard which reduce the amount of data. However, the results gave the idea of an important increment of the flexural strength due to the plaster. The results variation was higher for the specimens of the HCHB150P10 with a C.o.V. of 17.64% and lower for the group HCHB110 with a C.o.V. of 4.26%. Finally, Table 4.18 summarizes the flexural strength statistical parameters collected from the tests in agreement with the standard EN 1052-2 [16]. However, even considering that the tests were performed according to the standard, namely in terms of specimens’ geometric dimensions, and loading protocol,

116

4 Experimental Characterization of the As-Built Masonry …

(a)

(b)

(c)

(d)

(e) Fig. 4.25 Flexural strength tests parallel to horizontal bed joints: Flexural strength versus OOP displacement: a HCB110; b HCB150; c HCHB150P10; d HCB150 (existing) and HCHB150P10 (existing) and e global results

based on the observation of the test (type of failure), some considerations should be made regarding the results herein expressed. These considerations are related to the fact that these specimens were built with masonry units with certain percentage of voids (around 60%) which, combined with their reduced slenderness ratio, could have affected the results and not characterized well the flexural behaviour of infill panels made with this type of masonry units. For example, the specimens that were excluded from the set of tests presented in this work have evidenced shear failure

4.3 Stage 2—Mechanical Characterization of the Infill Masonry Wallets

117

Table 4.18 Flexural strength parallel to horizontal bed joints: statistical parameters Statistical parameter

HCHB110 HCHB150 HCHB150P100 HCHB150 HCHB150P10 (existing) (existing)

fb, mean, paralell (MPa) 0.12

0.14*

0.22

0.04

0.17

SD (M Pa)

0.01

0.02

0.04

N/A

N/A

C.o.V. (%)

4.26

12.63

17.64

N/A

N/A

fb,k,paralell (MPa)

0.08

0.15

N/A

N/A

0.09*

*

Number of tests with rupture according to the standard: 3 N/A—Not applicable

Fig. 4.26 Flexural strength tests parallel to horizontal bed joints: example of shear failure of HCHB150 specimen

outside the region between the OOP loading alignments. All these factors could have had an important effect in the test results. Due to this, additional tests must be carried out in long spandrels’ specimens (ideally with height equal to one storey height), even without any code standard framework. The comparison with such results will help to understand if the tests carried out according to minimum geometric dimensions required by the standard EN 1052-2 [16] are representative or not of the real expectable behaviour and capacity of panels made with these masonry units. Observed Damages From the observation of the specimens’ failure, all of them were characterized by the detachment between the masonry units and the horizontal bed joint, along the existing joints between the internal OOP loading application alignments distance. The specimens ruptured due to shear failure along the alignment of the loading application, as shown in Fig. 4.26, were not according to the standard. Figure 4.27 presents the most representative failure modes observed in each group of specimens.

118

4 Experimental Characterization of the As-Built Masonry …

Fig. 4.27 Flexural strength tests parallel to horizontal bed joints: Failure modes a HCHB110; bHCHB150; c HCHB150P10 (existing); and d HCHB150 (existing)

4.3.5.2

a)

b)

c)

d)

Flexural Strength Tests Perpendicular to the Horizontal Bed Joints

Complementarily, flexural strength tests perpendicular to the horizontal bed joints were carried out according to the standard EN1052-2 [16]. The standard NP EN10522 [16] prescribes the following specifications for the definition of the specimens’ dimensions (Fig. 4.28), namely: For masonry units with a height larger than 250 mm: – specimen width must have to be larger than 1000 mm; – at least one vertical joint should exist along the distance l2 . For masonry units with height equal or lower than 250 mm: – specimen width must have to respect these conditions: b ≥ 240 mm and b ≥ 3hu ; – at least one vertical and one horizontal joint should exist along the distance l2 . The adopted test setup was similar to the one described in the previous sub-section with the main differences related to the orientation of the OOP linear loading that,

4.3 Stage 2—Mechanical Characterization of the Infill Masonry Wallets

119

Fig. 4.28 Flexural strength tests perpendicular to horizontal bed joints: specimens’ dimensions according to NP EN1052-2 standard [16]

in this case is applied along the vertical direction as well as the restrain supports (Fig. 4.29). The specimens’ geometric dimensions were defined according to the standard requirements, namely 600 × 600 mm (width x height) with the load application distanced 300 mm from each other and the restrains 500 mm, respectively (Fig. 4.30). The instrumentation adopted for these tests was the same described in Sect. 5.3.5.1 and is illustrated in Fig. 4.30b. Similarly to the flexural strength tests parallel to the horizontal bed joints, the standard EN 1052-2 [16] only considers the tests in which the specimens’ rupture occurs between the OOP linear loading alignments.

Experimental Results Stress-Strain Curves The flexural strength was calculated according to Eq. 4.22 as the standard prescribes. From this set of tests, only one specimen from the HCHB110 group did not reach the rupture according to the standard; due to that, it was not considered for this study. From the force-displacement curves (Fig. 4.31), the following observations can be drawn: • The HCHB150 (existing) group reached the lowest flexural strength of the testing campaign (Fig. 4.31d), namely 19% lower than the specimens of the group HCHB150 and 5% than HCHB110 (Fig. 4.31a, b, respectively);

120

4 Experimental Characterization of the As-Built Masonry …

a)

b)

c) Fig. 4.29 Flexural strength tests perpendicular to horizontal bed joints: test setup a lateral view; b back view and c top view

a)

b)

Fig. 4.30 Flexural strength tests perpendicular to horizontal bed joints: a detail view and b instrumentation and specimens’ geometric dimensions

4.3 Stage 2—Mechanical Characterization of the Infill Masonry Wallets 0.5

0.3

0.2

HCHB110 Group fb,perpendicular,mean=0.27MPa

0.1

Specimen 1 Specimen 2 Specimen 3 Specimen 4 Specimen 5 μ curve μ−2SD curve μ+2SD curve

0.4

fb,perpendicular(MPa)

fb,perpendicular(MPa)

0.5

Specimen 1 Specimen 2 Specimen 3 Specimen 4 μ curve μ+2SD curve μ−2SD curve

0.4

121

0.3

0.2

HCHB150 Group fb,perpendicular,mean=0.32MPa

0.1

SD=0.06MPa CoV=18.10%

SD=0.08MPa C.o.V.=30.70% 0.0

0.0 0

1

2

3

4

5

6

7

0

1

2

OOP displacementaverage (mm)

a)

4

5

6

7

b) 0.5

0.5

Specimen 1 Specimen 2 Specimen 3 Specimen 4 Specimen 5 μ curve μ−2SD curve μ+2SD curve

0.3

0.2

HCHB150P10 Group fb,perpendicular,mean=0.30MPa

0.1

Specimen 1 Specimen 2 Specimen 3 μ curve μ−2SD curve μ+2SD curve

0.4

fb,perpendicular(MPa)

0.4

fb,perpendicular(MPa)

3

OOP displacementaverage (mm)

0.3

0.2

HCHB150 (Existing) Group fb,perpendicular,mean=0.26MPa

0.1

SD=0.07MPa CoV=28.40%

SD=0.02MPa CoV=7.90% 0.0

0.0 0

1

2

3

4

5

6

0

7

1

2

3

4

5

6

7

OOP displacementaverage (mm)

OOP displacementaverage (mm)

c)

d) 0.5

fb,perpendicular(MPa)

0.4

HCHB110 HCHB150 HCHB150P10 HCHB150 (Existing)

0.3

0.2

0.1

0.0 0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

OOP displacementaverage (mm)

e)

Fig. 4.31 Flexural strength tests perpendicular to horizontal bed joints stress-OOP displacement: a HCHB110; b HCHB150; c HCHB150P10; d HCHB150 (existing) and e global results

• The HCHB150 specimens achieved the highest flexural strength with a mean value of 0.322 MPa, 6% higher than the results obtained by the group HCHB150P10. The reason behind this difference can be attributed to the variability of the results; • Lower differences were observed when the specimens HCHB110 and the HCHB150 results were compared, with the first ones reaching an average value of 0.271 MPa, 11% less than the last. The HCHB150 (existing) group results are 19% lower than the ones reached by the as-built specimens. The group HCHB110 assumed the largest results’ variation

122

4 Experimental Characterization of the As-Built Masonry …

Table 4.19 Flexural strength perpendicular to horizontal bed joints: summary of test results Statistical parameter fb, mean, perpendicular (MPa)

HCHB110 0.27

HCHB150

HCHB150P10

HCHB150 (existing)

0.32

0.30

0.26

SD (M Pa)

0.08

0.06

0.02

0.07

C.o.V. (%)

30.30

18.10

7.90

28.40

0.18

0.22

0.20

N/A

fb,k,perpendicular (MPa)

with a C.o.V. equal to 30.3% and the lowest one by the group HCHB150P10 with C.o.V. equal to 7.9. Table 4.19 provides a summary of the flexural strength statistical parameters collected from the tests in agreement with the EN 1052-2 [16]. Damages Observed The typical failure modes occurred along these flexural tests were characterized by a pure vertical crack in the middle alignment of the specimens’ width. It was also observed total rupture of the central masonry units. Examples of the most representative failures observed in each group are presented in Fig. 4.32.

4.3.6 Final Remarks on Stage 2 Tests An experimental campaign was carried out aiming at characterizing the mechanical properties of infill masonry walls made with two HCHB units with two different thicknesses (110 mm and 150 mm). Additionally, a comparison between the results of as-built specimens and those collected from existing walls (built in the lab for full-scale tests) was performed to assess and discuss possible differences among them. Table 4.20 summarizes the main mechanical properties obtained within this experimental campaign. From the compression strength tests, the following conclusions can be highlighted: – The highest mean compressive strength and elastic modulus were obtained in the HCHB150 group, namely 40% and 8% respectively higher than HCHB110 group; – Fragile behaviour was observed in the wallets made with clay bricks with horizontal holes, without visible propagation of cracks. The collapse of parts of bricks with continuous increase of the vertical compressive force, increased the instability of the panel and resulted in fragile collapses of the specimens; – There is a need of additional tests to build a database with minimum number of data to revise and update analytical equations that correlate the infill panels’ compressive strength and their elastic modulus, since the current Eurocode 6 proposal seems to present low accuracy. From the diagonal tensile strength tests, the following conclusions can be highlighted:

4.3 Stage 2—Mechanical Characterization of the Infill Masonry Wallets

a)

b)

c)

d)

123

Fig. 4.32 Flexural strength tests perpendicular to horizontal bed joints failure modes: a HCHB110; b HCHB150; c HCB150P10; d HCHB150P10 (existing)

Table 4.20 Summary of mechanical properties obtained from the experimental campaign Mechanical Properties

HCHB110

HCHB150

HCHB150P10

fc,mean (MPa)

0.66

0.81

N/A

Emean (MPa)

1837

1975

N/A

Ss (MPa)

0.57

0.65

N/A

G (MPa)

1141

996

N/A

fb,parallel,mean (MPa)

0.12

0.14

0.22

fb,perpendicular,mean (MPa)

0.27

0.32

0.30

– Higher mean diagonal tensile strength and transverse stiffness were obtained in the HCHB150 group while the lower ones were achieved by the HCHB150 (existing) group; – The diagonal tensile strength obtained by the specimens HCHB150 was 14% higher than that obtained by HCHB110;

124

4 Experimental Characterization of the As-Built Masonry …

– The ratio transverse stiffness/elastic modulus was about 62% and 51% for the specimens with 110 and 150 mm thickness, respectively; – Fragile behaviour was observed in the wallets made with clay bricks with horizontal holes and some sliding failures. From the flexural strength tests, the following conclusions can be highlighted: – HCHB150 and HCHB110 specimens achieved the highest and lowest strength results, respectively; – Higher results variation was observed in the flexural strength tests perpendicular to the horizontal bed joints; – The plaster increased the flexural strength parallel to the horizontal bed joints of the specimens made with clay bricks with horizontal perforation around 60%; – The mortar adhesion properties play an important role in the flexural capacity of the specimens parallel to the horizontal bed joints. It would be useful to perform additional tests to complete this experimental campaign with additional compressive and diagonal tensile strength tests of HCHB150P10 and compressive strength, diagonal tensile strength and flexural strength tests of HCHB110P10 to clarify the plaster effect in the mechanical properties of infill wallets. Since the flexural strength tests were partially affected by shear failure mechanisms, as a future works the execution of flexural strength tests in specimens with larger dimensions (shear aspect ratio) and the corresponding results comparison would be important to validate the results herein presented.

4.4 Stage 3—Dynamic Characterization of Infill Panels Through Ambient Vibration Tests 4.4.1 Introduction The third stage of this experimental campaign is related to the ambient vibration tests carried out in infill panels to capture their OOP frequencies and corresponding vibration modes. Two different testing campaigns were made in laboratory and in-situ. The ambient vibration tests carried out in laboratory aimed to achieve the following goals: – Evaluation of the IP and OOP frequencies of a full-scale masonry infilled RC frame and the corresponding vibration modes; – Assessment of the panel natural frequencies’ variation during the postconstruction period; – Assessment of the impact of different level of axial load applied in the adjacent columns levels in the panels OOP frequencies; – Assessment of the impact of damage on the panels OOP frequencies.

4.4 Stage 3—Dynamic Characterization of Infill Panels Through Ambient …

125

The in-situ tests were carried out in three different buildings, where ambient vibration tests were performed in infill panels with different characteristics. The objectives of this testing campaign were: – Study of the panels’ geometric dimensions effect their OOP frequencies; – Study of the panels’ boundary conditions’ effect in their OOP frequencies; – Study of the openings’ effect in the panel OOP frequencies. In the present sub-section, the main test results and findings will be presented and discussed as well as the details regarding each infill panel studied; the testing methodology and setups adopted will be described.

4.4.2 Methodology The use of ambient vibration tests aims to study the infill walls dynamic characteristics and to understand the effect of some parameters: geometric dimensions, boundary conditions, masonry units, among others. From the literature review, very few studies can be found e.g. [24, 25], which highlights the relevance of this study and the need to understand the dynamic behaviour of the panel, thus possibly gathering information to build a database related to this topic. First, it is essential to instrument the infill panels with a sufficient number of accelerometers that allow capturing parameters such as OOP vibration frequencies, modes and/or damping. No information regarding this type of experimental tests is easily found in the literature due to several reasons: (i) lack of information regarding the infill panel’s dynamic characteristics; (ii) no access provided by the buildings’ owners to their interior for tests execution; and finally (iii) the requirement of studying the original infill panel behaviour without imparting any type of damage. To perform these ambient vibration tests it were used four different components: (i) a set of accelerometers PCB Piezotronics model 393B31 (0.5 g, 10.0 V/g, frequency range of 0.01-200 Hz); (ii) coaxial cables Piezotronics model 024R10; (iii) devices NI USB-9162 from National Instruments [26] that allow acquiring data results of all accelerometers simultaneously and finally, (iv) a portable computer to control the data acquisition (DAQ) and store results, as illustrated in Fig. 4.33. The methodology adopted to perform the ambient vibration tests is based on the following five procedures: (1) Definition of the test setup, namely the number and layout of the accelerometers. This decision is limited by the infill panel geometry, presence of openings and the panel’ boundary conditions (infill along borders). Several preliminary tests were made in laboratory to evaluate the minimum number of accelerometers required to capture the 1st OOP vibration frequency of an infill panel, constrained at four borders and without openings. From those tests, it was concluded that a minimum of 5 accelerometers is enough, with the distribution shown in Fig. 4.33a, b;

126

4 Experimental Characterization of the As-Built Masonry …

1 – Accelerometer 2 – Coaxial cable 3 - PC 4- Data acquisition device (NI USB 9162)

a)

b)

c)

0.04 0.03

Acceleration (g)

0.02 0.01 0.00 -0.01 -0.02 -0.03 -0.04 0

100

200

300

400

500

600

700

800

900

Time (s)

d)

Fig. 4.33 Ambient vibration tests in infill walls methodology: a schematic layout of the test setup; b general view of the test setup in the laboratory; c acquisition device; and d example of the evolution of the acceleration over time

4.4 Stage 3—Dynamic Characterization of Infill Panels Through Ambient …

127

(2) Placement of accelerometers on the infill panel using small steel plates that are fixed on it. This fixation to the infill panel can be made mechanically or using suitable glue; (3) Connection of accelerometers to the DAQ devices through coaxial cables (Fig. 4.33b); (4) Connection between the DAQ devices and the PC through USB cables (Fig. 4.33c); (5) Make the data acquisition, composed of three sets of measurements, during periods of 15 min each (Fig. 4.33d), with a sampling frequency equal to 2000 Hz. The modal analysis was carried out through the enhanced frequency domain decomposition (EFDD) technique (Fig. 4.34a), implemented in the ARTeMIS Extractor software [27], from which the natural frequencies and the corresponding vibration modes were captured. More details regarding this technique can be found in [28]. It is observed that the first OOP vibration mode of an infill masonry panel is characterized by mobilizing the almost whole panel in the same OOP direction simultaneously, evidenced in the centre of the panel, as observed in Fig. 4.34b. Although it is possible to find different symmetric and asymmetric OOP modes (Fig. 4.34c, d), this work mainly focuses on the 1st OOP vibration mode. Regarding the measurement of the first in-plane (IP) natural frequency of the infill panel, it must be considered that this frequency is more influenced by the contribution of the surrounding RC frame since they interact directly with each other. The typical 1st IP modal configuration of an infilled RC frame is plotted in Fig. 4.34c and is characterized by a progressive deformation from the bottom until the top of the frame, being the largest deformation in the top of the frames height. Note that these considerations only refer to infill panels with four constrained borders.

4.4.3 Laboratory Tests 4.4.3.1

Test Specimens and Setup

The laboratory tests were carried out at the Laboratory for Earthquake and Structural Engineering (LESE) in a full-scale, one storey, one bay infilled RC frame that was built in the scope of the experimental characterization of the OOP capacity of infill panels that will be later presented in Chap. 6. The general dimensions of the specimen (RC frame included) are 4.80 × 3.30 m (width x height) and the cross sections of the RC columns and beams are 0.30 × 0.30 m and 0.30 × 0.50 m, respectively. Horizontal hollow clay bricks, 150 mm thick, were used to construct a 2.30 × 4.20 m infill panel. No reinforcement was used to connect the infill panel and the surrounding RC frame. The tests were carried out in a total of 6 infill panels prior to their destructive testing (the IP and OOP tests will be later described in Chap. 5). The first objective of this study is to characterize the OOP and IP natural frequencies and vibration modes.

128

4 Experimental Characterization of the As-Built Masonry …

a)

c)

b)

d)

e)

f)

Fig. 4.34 Examples of modal identification test results: a peak picking of the singular values of spectral density matrices; b 1st OOP vibration mode (panel); c 2nd OOP vibration mode (panel); d 3rd OOP vibration mode (panel); e 1st IP vibration mode (global mode) and f 2nd IP vibration mode (global mode) (green—mode shape; blue—original position)

4.4 Stage 3—Dynamic Characterization of Infill Panels Through Ambient …

129

a)

b)

c)

Fig. 4.35 Modal identification test in LESE speciment: a layout of the test setup; b lateral view; and c front view

Secondly, ambient vibration tests were performed during the period after construction to assess the evolution of the IP and OOP frequencies over time. Finally, different gravity load levels were applied several days after panel construction at the top of the adjacent RC frame’ columns to assess their effect in the panel OOP vibration frequencies. For the whole set RC frame plus the infill panel the ambient vibration tests were carried out using twelve accelerometers that were distributed on the RC frame (7 accelerometers—4 for IP vibration measurement and 3 for OOP vibration measurement) and the infill panel (5 accelerometers—OOP vibration measurement) as illustrated in Fig. 4.35, to achieve the IP and OOP vibration frequencies.

4.4.3.2

Modal Identification of As-Built Infill Panels

Ambient vibration tests were carried out in as-built (i.e. non strengthened) infill panels with and without plaster to obtain the OOP and IP natural frequencies and

130

4 Experimental Characterization of the As-Built Masonry …

vibration modes of the specimen above described. The specimens tested to determine the OOP frequencies were: (1) without plaster Inf_02 and Inf_05; (2); with plaster Inf_06, Inf_08, Inf_09 and Inf_11. Additionally, it was included the average, standard deviation (SD) and coefficient of variation (C.o.V) for each typology (with and without plaster). The results are shown in Table 4.21, from which it is possible to observe that the average 1st OOP vibration mode frequency of the panels without plaster was 22.5 Hz, about 3% higher than the 1st OOP vibration mode frequency of the panels with plaster. The plaster increases the panel stiffness, however increases at the same time the panel mass which justifies the results concerning the 1st OOP vibration mode frequency. The 2nd OOP vibration mode of the panels without plaster was about 40.3 Hz, 11% lower than those with plaster. The 3rd OOP vibration mode frequency of the panels without plaster was 58.1 Hz, about 5% lower than the results obtained by the panels with plaster. These results are in the same range as De Angelis et al. [25]. Concerning the IP frequencies and vibration modes, two specimens were tested, one without plaster (Inf_05) and one with plaster (Inf_11). The obtained 1st and 2nd global IP vibration mode frequencies are presented in Table 4.22. However, the results herein shown have to be observed carefully, since both frames were damaged due to the panel OOP tests that were performed which caused some cracking, and spalling of some parts of the frame. This obviously affects the IP frequencies, since it reduces the frame’s stiffness. Due to that, these results are simply presented here, but no significant discussion and findings can be extracted from them. Table 4.21 Summary of the OOP frequencies obtained in the ambient vibration tests in the Laboratory Specimen

Plaster

Inf_02

No

Inf_05

Inf_06 Inf_08

Yes (10 mm)

Panel 1st OOP mode (Hz)

Panel 2nd OOP mode (Hz)

Panel 3rd OOP mode (Hz)

24.3

38.4

58.6

20.8

42.3

57.6

Average: 22.5 SD: 2.5 C.o.V.(%): 11.1

Average: 40.3 SD: 2.8 C.o.V.(%): 6.8

Average: 58.1 SD: 0.7 C.o.V.(%): 1.3

21.3

41.9

61.3

20.6

44.6

65.3

Inf_09

21.7

44.1

67.1

Inf_11

22.5

45.9

68.6

Average: 22.2 SD: 0.6 C.o.V.(%): 2.5

Average: 44.9 SD: 1.2 C.o.V.(%): 2.7

Average: 67.8 SD: 1.1 C.o.V.(%): 1.6

4.4 Stage 3—Dynamic Characterization of Infill Panels Through Ambient …

131

Table 4.22 Summary of the IP frequencies obtained in the ambient vibration tests in the laboratory Specimen

Plaster

Global first IP vibration mode (Hz)

Global second IP vibration mode (Hz)

Inf_05

No

17.20

54.36

Inf_11

Yes (10 mm)

14.85

40.08

4.4.3.3

Columns’ Axial Load Effect in the Panel OOP Frequencies

This stage of ambient vibration tests carried out in the laboratory aimed to assess the effect of applying different levels of axial load in the RC frame’ columns in the panel’ OOP frequencies. For this, different axial load levels were applied in each column, namely 30k N (ν = 0.02); 85k N (ν = 0.05); 150 kN (ν = 0.08); 221 kN (ν = 0.12) and 273 kN (ν = 0.14). Ambient vibration tests were performed prior and after this set of tests to evaluate any possible effect of the columns axial load application in the panel natural frequencies. Table 4.23 summarizes the panel OOP frequencies obtained for each axial load level. From the results, it was found that the natural frequencies increased about 17% until reaching N = 273 kN. Since the infill panel mass did not suffered any variation during the tests, it was possible to compute the infill panel OOP stiffness variation. Thus, for the maximum axial load level applied (N = 273 kN), the panel OOP stiffness increased around 38% compared with the initial situation without axial load (i.e. N = 0 kN). To conclude, it is also important to say that additional tests were performed after the infill panel construction along six different moments, namely: (i) wall construction day (Day 0); (ii) 1 day after construction (Day 1); (iii) 4 days after construction; (iv) 25 days after construction (Day 25) and (v) 42 days after construction (Day 42). The results showed that no significant variation occurred for the period larger than 4 days after construction. Table 4.23 OOP frequencies evolution for different axial load levels t (days)

Level

42

0

N (kN)

f (Hz)

C.o.V. (%)(a)

0

19.28

3.96

1

30

20.01

0.33

2

85

21.34

1.06

3

150

22.35

0.33

4

221

22.31

0.30

5

273

22.64

1.01

132

4 Experimental Characterization of the As-Built Masonry …

4.4.4 In-Situ Tests This sub-section presents an experimental campaign of ambient vibration tests performed in-situ. The architectural and structural description of each of the three buildings where the tests were performed will be firstly detailed as well as the characteristics of the tested infill panels. Afterwards, the global results will be presented and discussed in terms of assessment of the effect of parameters such as geometric dimensions, boundary conditions, slenderness and openings.

4.4.4.1

Building A

The building A is a residential one, under construction by the time of the tests, and located in Porto (Portugal). The structure has an in-plan rectangular geometry, as shown in Fig. 4.36, with a ground floor area of 485 m2 and a total height of 22.5 m (four storeys plus the ground floor). The building structure comprises 6 and 4 alignments in the transverse and longitudinal direction, respectively. HCHB units with different thicknesses (110 mm, 150 mm and 220 mm) composed the external and internal infill walls. The geometric dimensions of each type of brick are shown in Fig. 4.37. This building is characterized by a vertical irregularity regarding the distribution of the IM walls, with reduced number of internal partition walls exists in the groundfloor to be used for commercial purposes. It was observed that some external partition walls were built partially supported (only in 3/4 of the 220 mm brick thickness) on the bottom beams or slabs in order to account for thermal bridges’ correction purposes (Fig. 4.38a). Regarding the internal partition walls, some were composed by one-leaf panels 110 mm thick (Fig. 4.38b) and others of double-leaf panels, 110 + 110 mm plus the gap between the leaves filled with expanded polystyrene plates, as shown in Fig. 4.38c. From building A, fourteen infill walls with different geometric dimensions (length and height), different thicknesses, openings and boundary conditions were tested, all of them made of HCHB units. Some of the tested panels had holes/cavities drilled for electrical instalations (walls A11, A12 and A13) which could pottentially affect the test results. It is also important to say that the height of the walls A1, A2 and A3 are lower when compared with remaining ones. In fact, those panels are internal partition walls that were defined to have those dimensions due to architectural reasons. Table 4.24 provides a summary of the main infill walls’ characteristics, whose geometry is shown in Fig. 4.39. The data acquisition of the acceleration measurements obtained with the accelerometers was performed using the LabVIEW SignalExpress software [29]. The modal analysis of the specimens was performed from which the first natural frequency of every single panel was obtained; the results are summarized in Table 4.25. For a better understanding of the results, different variables were considered: the

4.4 Stage 3—Dynamic Characterization of Infill Panels Through Ambient …

133

Fig. 4.36 Building A: a in-plan layout; and b general elevation view

300x200x110mm

300x200x150mm

300x200x220mm

a)

b)

c)

Fig. 4.37 Masonry units geometric dimensions (mm) a HCHB110; b HCHB150 and c HCHB220

134

4 Experimental Characterization of the As-Built Masonry …

b)

a)

c)

Fig. 4.38 Building A—Infill panels general view: a external wall partially supported; b elevation view; and c cross-section view

Table 4.24 Building A: Tested infill walls characteristics Wall

L (m)

H (m)

Thickness (m)

Opening

A1

3.50

1.30

0.11

No

A2

3.50

1.30

0.11

A3

3.55

3.20

A4

3.20

A5

Panel area (m2 )

Openings’ area Aopening (m2 )

Floor location

4.55

–

GF(a)

No

4.55

–

GF(a)

0.11

No

11.36

–

GF(a)

0.95

0.11

No

3.04

–

GF(a)

3.55

3.20

0.11

No

11.36

–

GF(a)

A6

2.75

3.20

0.11

Door

8.8

2.10

GF(a)

A7

2.45

3.20

0.11

Door

8.8

2.10

GF(a)

A8

5.00

1.75

0.15

No

8.75

–

4th

A9

2.75

2.30

0.15

No

6.33

–

4th

A10

3.60

1.00

0.11

No

3.6

–

3rd

A11

3.20

2.80

0.22

Window

8.96

3.05

3rd

A12

3.40

2.80

0.22

No

9.52

–

3rd

A13

3.70

2.80

0.22

No

10.36

–

3rd

A14

2.10

3.20

0.11

No

6.72

–

3rd

GF– Ground-floor

infill walls aspect ratio (length/height ratio), the infill panel diagonal length (D), the openings area through the parameter μ (opening ratio), given by Eq. 4.23, boundary conditions (BC—Border constrains; 4BC—4 borders constrained; 3BC—3 borders constrained) and panel thickness (t). Table 4.25 is sorted according to the boundary conditions group (4BC or 3BC) and the presence of openings for a better comparison among the results herein presented.

4.4 Stage 3—Dynamic Characterization of Infill Panels Through Ambient …

( μ=1−

Aopening Linf × Hinf

135

) (4.23)

where Aopening is the opening area and Linf and Hinf are the infill panel length and height, respectively. The parameter μ is equal to 1 when the panel has no opening. From the analysis of the 1st OOP frequency of each panel summarized in Table 4.25 and plotted in Fig. 4.40, it can be found that the OOP frequencies of the panels without openings, with four borders constrained (4BC) varied between 24.5 Hz and 34.2 Hz; concerning the panels with three borders constrained (3BC) the variation is between 14.7 Hz and 22.4 Hz. The results obtained by the panels with openings and four borders constrained (A6, A7 and A11) ranged between11.5 Hz and 22.7 Hz.

A1

A2 Fig. 4.39 Building A—tested infill walls geometric dimensions

136

4 Experimental Characterization of the As-Built Masonry …

A3

A4 Fig. 4.39 (continued)

Through the analysis of the results obtained by the panels 110 mm thick, it is observed again a large variation between 15 to 35 Hz. For the panels tested with the thickness equal to 150 mm, the OOP frequencies ranged from 17 and 35 Hz and finally, the panels made with masonry units 220 mm thick, the results varied between 12 to 30 Hz. Finally, it is also important to mention that for infill panels with 4 borders constrained (4BC), the frequencies are 20–60% higher than the ones obtained by panels with 3 borders constrained (3BC). The results regarding the infill walls aspect ratio (length/height ratio) are presented in Fig. 4.41a, from which is possible to observe a large variation of the results that were divided according to the panel thickness. The infill panels’ length/height ratio

4.4 Stage 3—Dynamic Characterization of Infill Panels Through Ambient …

A5

A6

Fig. 4.39 (continued)

137

138

4 Experimental Characterization of the As-Built Masonry …

A7

A8 Fig. 4.39 (continued)

of the panels 110 mm thick varies between 0.7 and 3.6 and the variation of the OOP frequencies is around 14.7 Hz to 34.1 Hz. The infill panels’ length/height ratio of the panels 150 mm thick varies between 1.2 and 2.9 and the variation of the OOP frequencies is around 18 Hz to 34.2 Hz. Finally, the infill panels’ Linf /Hinf ratio of the panels 220 mm thick varies between 1.14 and 1.32 and the variation of the OOP frequencies is around 11.5 Hz to 28.2 Hz. For example, panels A1 and A2 have the same relation Linf/Hinf, without openings and same border constraints; however, the results are 25% different, probably due to the workmanship or experimental variability that can affect the panel OOP behaviour. For a better analysis of the Linf /Hinf ratio effect, the global results (Fig. 4.41a) were divided into three groups: wall panels made with HCHB110 units (Fig. 4.41b); HCHB150 units (Fig. 4.41c) and HCHB220 (Fig. 4.41d). Regarding the wall panels

4.4 Stage 3—Dynamic Characterization of Infill Panels Through Ambient …

139

A9

A10 Fig. 4.39 (continued)

made with HCHB110 units, a slight trend of increase of the OOP vibration frequencies with the increase of the Linf/Hinf ratio can be found. Even the panels 110 mm thick with openings (only 3 specimens) present an increase of the OOP frequencies with the increase of the Linf /Hinf ratio. Due to the reduced number of tests for the panels made with HCHB150 and HCHB220 units, it is impossible to draw any conclusion regarding the Linf /Hinf ratio effect.

140

4 Experimental Characterization of the As-Built Masonry …

A11

A12 Fig. 4.39 (continued)

4.4 Stage 3—Dynamic Characterization of Infill Panels Through Ambient …

A13

A14 Fig. 4.39 (continued)

141

142

4 Experimental Characterization of the As-Built Masonry …

Table 4.25 Building A: Ambient vibration test results Wall

Linf /Hinf

D (m)

μ

Opening

Border constraints

t (m)

1st OOP Frequency (Hz)

A1

2.69

3.73

1

No

4 BC

0.11

31.7

A2

2.69

3.73

1

No

4 BC

0.11

24.5

A5

1.10

4.78

1

No

4 BC

0.11

30.0

A9

1.20

5.59

1

No

4 BC

0.15

34.2

A10

3.60

3.74

1

No

4 BC

0.11

34.1

A12

1.21

4.40

1

No

4 BC

0.22

28.2

A13

1.32

4.64

1

No

4 BC

0.22

27.6

A6

0.86

4.22

0.76

Door

4 BC

0.11

22.5

A7

0.77

4.03

0.73

Door

4 BC

0.11

22.7

A11

1.14

4.25

0.66

Window

4 BC

0.22

11.5

A3

1.11

4.78

1

No

3 BC

0.11

16.5

A4

3.48

3.33

1

No

3 BC

0.11

22.4

A8

2.86

5.30

1

No

3 BC

0.15

18.0

A14

0.66

3.83

1

No

3 BC

0.11

14.7

40

Building A

t=0.11m t=0.15m t=0.22m

1st Out-of-plane frequency (Hz)

4BC 4BC 4BC 4BC

30

4BC 4BC 4BC 3BC

4BC 4BC

20

3BC 3BC Door Door

3BC 4BC Window

10

0

0 A1 A 2 A 3 A4 A5 A 6 A7 A8 A9 A10 A11 A12 A13 A14 15

Wall Panels Fig. 4.40 Building A test results: 1st OOP frequency

The same strategy was adopted to analyse the effect of the diagonal dimension D. The global results (plotted in Fig. 4.42a) were divided according to each wall panel wall thickness group. Again, despite the larger number of tests performed in panels

4.4 Stage 3—Dynamic Characterization of Infill Panels Through Ambient … 40

40

Building A

35

30 25 20 15 10

t=0.11m t=0.15m t=0.22m

5 0 0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

1st Out-of-plane frequency (Hz)

1st Out-of-plane frequency (Hz)

35

Building A t=0.11m

30 25 20 15 10

3BC Without opening With opening

5 0 0.0

4.0

0.5

1.0

1.5

Linf/Hinf ratio

2.5

35

30 25 20 15 10 5

0.5

1.0

1.5

2.0

Linf/Hinf ratio

c)

3.5

4.0

Building A t=0.22m

30 25 20 15 10 5

without openings with openings

3BC 0 0.0

3.0

b) 40

Building A t=0.15m

1st Out-of-plane frequency (Hz)

1st Out-of-plane frequency (Hz)

35

2.0

Linf/Hinf ratio

a) 40

143

2.5

3.0

3.5

4.0

0 0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

Linf/Hinf ratio

d)

Fig. 4.41 Building A test results: Linf /Hinf ratio: a global results; b panels 110 mm thick; c panels 150 mm thick, and d panels 220 mm thick

made with HCHB110 units (Fig. 4.42b), it is impossible to observe any trend due to the dispersion of the results regarding the panels such as with and without openings. Also, the results of the panels with 3BC are not enough to support any conclusion regarding this effect. For example, for a diagonal dimension D equal to 3.33 m the OOP vibration frequency is 22.4 Hz, which is 35% higher than the result obtained by the wall panel with 4.78 m. Globally, regarding the panels 3BC without openings it can be found that the OOP vibration frequencies decreased with the increase of D. Again, due to the reduced number of tests in panels made with HCHB150 and HCHB220 units (Fig. 4.42c, d) it is not possible to observe any trend. Finally, no clear evidence of a relationship is observed between the openings ratio μ and the infill panels’ OOP frequencies because the amount of data is still scarce (Fig. 4.43). However, as expected, it is observed that the OOP frequency is lower for panels with openings. To conclude, and due to the dispersion of results and absence of visible trends, it must be stated that the OOP frequencies of the panels without openings and 4BC varied between 24.5 Hz and 34.2 Hz; concerning the panels without openings and

144

4 Experimental Characterization of the As-Built Masonry …

1st Out-of-plane frequency (Hz)

40

Building A t=0.11m t=0.15m t=0.22m

35 30

35

1st Out-of-plane frequency (Hz)

40

25 20 15 10 5 0 2.0

2.5

3.0

3.5

4.0

4.5

5.0

5.5

Building A t=0.11m

30 25 20 15 10

3BC Without opening With opening

5 0 2.0

6.0

2.5

3.0

3.5

a)

1st Out-of-plane frequency (Hz)

35

4.5

Building A t=0.15m

35

30 25 20 15 10 5

2.5

3.0

3.5

5.5

6.0

Building A t=0.22m

30 25 20 15 10 5

Without openings With openings

3BC 0 2.0

5.0

b) 40

1st Out-of-plane frequency (Hz)

40

4.0

Diagonal dimension D (m)

Diagonal dimension D (m)

4.0

4.5

5.0

5.5

0 2.0

6.0

2.5

Diagonal dimension D (m)

3.0

3.5

4.0

4.5

5.0

5.5

6.0

Diagonal dimension D (m)

c)

d)

Fig. 4.42 Building A test results: Diagonal dimension D: a global results; b panels 110 mm thick; c panels 150 mm thick, and d panels 220 mm thick 40

Building A

1st Out-of-plane frequency (Hz)

35 30

t=0.11m t=0.15m t=0.22m

25 20 15 10 5 0 0.5

0.6

0.7

0.8

μ Fig. 4.43 Building A test results: OOP frequency versus μ

0.9

1.0

1.1

4.4 Stage 3—Dynamic Characterization of Infill Panels Through Ambient …

145

3BC the variation was from 14.7 Hz to 22.4 Hz. The panels with openings and 4BC obtained results from 11.5 Hz to 22.7 Hz.

4.4.4.2

Building B

The second group of tests was conducted on a residential building (Building B), in Gaia (Portugal). The building has a regular geometry with a ground-floor area of 275 m2 , four upper-storeys including the ground-floor, which has a total height of 16 m. The building structure is composed by 6 transverse and 4 longitudinal alignments, as illustrated in Fig. 4.44. The infill panels in this building are made with HCHB110 units. Different typologies were found, namely panels without openings, panels with windows, panels with doors or panels with combined window plus door. A slight increase of the openings’ size was observed in the ground-floor walls. Part of the external partition walls was composed of one-leaf infill walls 110 mm thick and the remaining ones of double-leaf panels 110 + 110 mm with the gap between leaves filled with expanded polystyrene plates. Three external infill panels located in the 3rd storey, with the same thickness, geometry and boundary conditions were tested in Building B. The geometry of the panels tested and the layout of the accelerometers adopted for the ambient vibration tests are shown in Fig. 4.45. Table 4.26 summarizes each tested wall’s major characteristics and geometric dimensions. Table 4.27 presents the experimental OOP vibration frequencies obtained for each tested infill wall. From the test results, it can be found that the wall panel B3 achieved the highest OOP frequency, which can be explained by the absence of openings and because it has a shorter length than the wall panel B1, which increased its OOP stiffness, as observed already in Building A (Fig. 4.46a). From the plot OOP frequencies according to the H/L ratio (Fig. 4.46b) it can be observed that the higher Linf /Hinf ratio of wall B1 is responsible by 20% lower

a)

b)

Fig. 4.44 Building B: a in-plan disposition; and b general elevation overview

146

4 Experimental Characterization of the As-Built Masonry …

Wall B1 a)

Wall B2 b) Fig. 4.45 Building B—Tested wall panels’ geometries: a Wall B1; b Wall B2 and c Wall B3

frequency than that of the wall B3. It can be observed that the OOP frequency of wall B2 is clearly affected by the existence of the opening, which reduced the panel frequency significantly. It is also observed, in Fig. 4.46c, that the increase of D leads to lower OOP frequencies, as proved by the wall B3. Finally, it can be also observed that the existence of openings (20% of the panel area) reduced the OOP frequency by about 20% (Fig. 4.46d).

4.4 Stage 3—Dynamic Characterization of Infill Panels Through Ambient …

147

Wall B3 c) Fig. 4.45 (continued)

Table 4.26 Building B: Characteristics of the tested wall panels Wall panel

Linf (m)

Hinf (m)

Thickness (m)

Opening

Panel area (m2 )

Openings’ area (m2 )

Floor location

B1

3.90

2.70

0.11

No

10.53

–

3rd

B2

3.80

2.70

0.11

Window

8.28

2.25

3rd

B3

3.50

2.70

0.11

No

9.45

–

3rd

Table 4.27 Building B: ambient vibration results Wall panel

Linf /Hinf

D (m)

μ

Opening

Border constraints

1st OOP Frequency (Hz)

B1

1.44

4.74

1

No

4BC

22.4

B2

1.40

4.66

0.79

Yes

4BC

18.2

B3

1.30

4.42

1

No

4BC

27.3

4.4.4.3

Building C

The third group of in-situ tests was carried out in Building C, located in Viseu (Portugal) and represents a typical construction of the 1960’s in Portugal (Fig. 4.47a).

148

4 Experimental Characterization of the As-Built Masonry … 40

Building B t=0.11m

30

20

window

10

t=0.11m

B2

B1

B3

30 25 20 15 10 5 0 0.0

0 0

Building B

35

1st Out-of-plane frequency (Hz)

1st Out-of-plane frequency (Hz)

40

4

0.2

0.4

0.6

a)

1st Out-of-plane frequency (Hz)

35

Building B t=0.11m

35

30 25 20 15 10 5 0 2.0

2.5

3.0

3.5

4.0

1.0

1.2

1.4

1.6

1.8

2.0

b) 40

1st Out-of-plane frequency (Hz)

40

0.8

Linf/Hinf

Wall panel

4.5

5.0

5.5

t=0.11m

30 25 20 15 10 5 0 0.5

6.0

Building B

0.6

0.7

0.8

0.9

1.0

1.1

μ

Diagonal dimension D (m)

c)

d)

Fig. 4.46 Building B test results: a 1st OOP frequency; b OOP frequency versus L/H ratio; c OOP frequency versus diagonal dimension D and d OOP frequency versus μ

The block plan is rectangular, 11.45 m wide, 17.40 m long (Fig. 4.47b, c), and the building has 13.30 m total height (excluding the basement height) comprising 4 living floors plus the ground-floor’ height. The main structural system (6 transverse parallel plane frames) characterizes the building layout, the distance between the frames’ axes being about 3.80 m.

a)

b)

Fig. 4.47 Building C a General view b Transversal view and c Building plan

c)

4.4 Stage 3—Dynamic Characterization of Infill Panels Through Ambient …

149

The first floor corresponds to a basement/garage without internal partition walls. Above this level there are four successive living floors; the entrance in the front facade is located at the second-floor level. The top floor is accessible and has small individual spaces for residents’ belongings. HCHB units make the infill walls with thicknesses 70 mm, 110 mm and 150 mm. Several typologies can be found in Building C: for example the external walls of front and rar façades have panels with openings (such as windows or doors). The lateral external walls are composed of double-leaf panels with different combinations such as 110 + 110 mm, 80 mm + 110 mm or 150 mm + 110 mm with the gap between the leaves filled with expanded polystyrene plates. The experimental campaign in Building C focused on three external infill walls, located in different storeys. The walls C1 and C2 are double-leaf IM walls with different characteristics: the wall C1 is composed of two leaves with 150 mm + 80 mm thickness and the wall C2 has 150 + 110 mm leaves’ thickness. Both have different types of openings as summarized in Table 4.5. These double-leaf walls were tested independently in both the interior (C1_int) and exterior panels (C1_ext) since no connections were enhanced. The wall C3 is a rectangular wall with no opening with 110 mm thickness. Figure 4.48 shows the general view of the walls C1_ext and C3 during the ambient vibration tests. Table 4.28 summarizes the main characteristics of each infill panel tested in Building C. For these tests, the adopted number of accelerometers increased for up to 20 accelerometers in the case of the panel without opening (Wall C3) and 8 accelerometers for the remaining panels (C1 and C2 int and ext). The geometry of each wall panel is illustrated in Fig. 4.49.

Experimental Results Table 4.29 presents the data results obtained for each infill panel. From the results (plotted in Fig. 4.50a), it can be observed that the panels with openings have similar

a)

b)

Fig. 4.48 Building C: General view of the tested infill walls a Panel C1_ext b Panel C3

150

4 Experimental Characterization of the As-Built Masonry …

Table 4.28 Building C: Tested IMW characteristics Wall

Linf (m)

Hinf (m)

Thickness (m)

Opening

Área of the panel (m2)

Openings area (m2)

Floor location

C1_int

3.15

2.55

0.07

Window

8.16

2.94

1st

C1_ext

3.15

2.55

0.15

Window

8.16

2.94

1st

C2_int

3.05

2.55

0.11

Door

7.78

4

3rd

C2_ext

3.05

2.55

0.15

Door

C3

3.50

2.80

0.11

No

7.78

4

3rd

11.36

–

3rd

OOP frequencies independent of the panel thickness, as proved by the difference of 0.2% between C1_int and C1_ext and 12% between C2_int and C2_ext. It can be observed that the double-leaf infill panels reached significantly higher OOP frequencies, about twice than the single leaf panel. This unexpected result could suggest the possible (and not identified) internal connection between the internal and external panel). Regarding the H/L ratio (plotted in Fig. 4.50b), for a similar H/L ratio there is a large scatter of results. For example, for Linf /Hinf around 1.25 there is a variation from 31 to 65 Hz depending on the panel thickness and openings percentage. Similar observation can be performed through the analysis of the plot of the diagonal dimension D effect, plotted in Fig. 4.50c. Finally it can be observed that the panel C3 has the lowest OOP frequency, even when compared with the panels with openings (Fig. 4.50d). The comparison between the panel C1_int and C1_ext shows that the results are similar with a difference smaller than 1%. Larger difference is found between the panel C2_int and C2_ext, which is around 13%. From the results of the double-leaf panels, the frequencies are significantly higher, for the same panel thickness, than the results obtained in the previous buildings studied in this campaign. This can be related to some contribution of the two panels for the OOP vibration frequencies, besides the abcense of visible mechanical connection between the leaves. These results reinforce the idea that the study of double-leaf walls should be performed separetly from the single leaf panels, since the behaviour appears to be different.

4.4.4.4

Global Results

All the results collected from the ambient vibration tests are now analysed together to allow global comparisons and better assess each parameter’s effect under analysis. Figure 4.51a presents the distribution of the OOP frequencies according to each group of panels and the ratio Linf /Hinf . For panels made with HCHB110 units (the group with a larger number of specimens) no significant variations can be observed in the panels without openings and the OOP frequencies range from 15 to 35 Hz (except in the case of the interior leaf of the double-leaf panel of building C).

4.4 Stage 3—Dynamic Characterization of Infill Panels Through Ambient …

151

Wall C1_int and Wall C1_ext a)

Wall C2_int and Wall C2_ext b) Fig. 4.49 Building C—geometry of the tested infill wall: a Panel C1; b Panel C2 and c Panel C3

152

4 Experimental Characterization of the As-Built Masonry …

Wall C3 c) Fig. 4.49 (continued)

Table 4.29 Building C: Ambient vibration test results Wall

Linf /Hinf

D (m)

μ

Opening

Border constraints

1st OOP frequency (Hz)

C1_int

1.24

4.05

0.63

Yes

4 BC

53.57

C1_ext

1.24

4.05

0.63

Yes

4 BC

53.48

C2_int

1.20

3.82

0.49

Yes

4 BC

64.45

C2_ext

1.20

3.82

0.49

Yes

4 BC

57.13

C3

1.25

4.42

1

No

4 BC

31.56

However, regarding the panels made with HHCB110 units with openings it is visible that for larger L/H ratio the OOP vibration frequency increases. Regarding the remaining groups (made with HCHB150 and HCHB220 units with and without openings); due to the reduced number of specimens, it is impossible to draw any conclusion regarding the L/H ratio effect. The influence of the diagonal dimension D in the OOP frequencies is plotted in Fig. 4.51b. Larger dispersion of results was found for the panels made with HCHB110 units with and without openings. Again, the workmanship and experimental variability can justify this variability. Lastly, the effect of the opening was

4.4 Stage 3—Dynamic Characterization of Infill Panels Through Ambient …

1st Out-of-plane frequency (Hz)

1st Out-of-plane frequency (Hz)

75

Building C

t=0.07m t=0.15m t=0.11m

60

40

20

Building C t=0.07m t=0.11m t=0.15m

60

45

30

15

Double-leaf wall C1_int and C1_ext Double-leaf wall C2_int and C2_ext 0 0.00

0

C1_int

C2_int

C1_ext

C3

C2_ext

0.25

0.50

0.75

a) 75

1st Out-of-plane frequency (Hz)

t=0.07m t=0.11m t=0.15m

45

30

15

Double-leaf wall C1_int and C1_ext

1.75

2.00

60

t=0.07m t=0.11m t=0.15m

45

30

15

Double-leaf wall C1_int and C1_ext Double-leaf wall C2_int and C2_ext

0 3

1.50

Building C

Double-leaf wall C2_int and C2_ext 2

1.25

b) 75

Building C

60

1.00

Linf/Hinf ratio

Wall number

1st Out-of-plane frequency (Hz)

153

4

5

6

0 0.0

0.2

0.4

0.6

c)

0.8

1.0

μ

Diagonal dimension D (m)

d)

Fig. 4.50 Building C test results: a 1st OOP frequency; b OOP frequency versus L/H ratio; c OOP frequency versus D and d OOP frequency versus μ

assessed (Fig. 4.51c). From the results, it must be stated again that further tests are needed to extract valid conclusions and to quantify their effect (total amount of tests equal to 8).

4.4.5 Final Remarks on Stage 3 Tests Ambient vibration tests were carried out in laboratory on a full-scale infilled RC frame, aiming at: (i) studying the evolution of the infill walls IP and OOP frequencies along the period after construction of the panel; and (ii) evaluation of the axial load (in the frame columns) effect in the panel OOP frequencies of the panel. These tests showed that the average 1st OOP vibration mode frequency of the panels without plaster was about 22.5 Hz, about 3% higher than the 1st OOP vibration mode frequency of the panels with plaster. It was found that besides the increment of the panel stiffness due to the plaster, there is an increment at the same time of the panel mass which justifies the results concerning the 1st OOP vibration mode

154

4 Experimental Characterization of the As-Built Masonry …

60

50 45 40

55

1st Out-of-plane frequency (Hz)

1st Out-of-plane frequency (Hz)

60

t=0.07m w/o openings t=0.11m with openings t=0.11m w/o openings t=0.15m with openings t=0.15m w/o openings t=0.22m with openings t=0.22m w/o openings Double-leaf walls

55

35 30 25 20 15 10 5 0 0.0

50 45 40 35 30 25

t=0.07m w/o openings t=0.11m with openings t=0.11m w/o openings t=0.15m with openings t=0.15m w/o openings t=0.22m with openings t=0.22m w/o openings Double-leaf walls

20 15 10 5 0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

4.5

0

5.0

1

2

3

4

5

6

Diagonal dimension D (m)

L/H ratio

a)

b)

1st Out-of-plane frequency (Hz)

75

60

t=0.07m t=0.11m t=0.15m t=0.22m Double-leaf walls

45

30

15

0 0.0

0.2

0.4

0.6

0.8

1.0

μ

c)

Fig. 4.51 Global results: a OOP frequency versus L/H ratio; b OOP frequency versus diagonal dimension D and c OOP frequency versus μ

frequency. Slightly larger variations were found for the 2nd and 3rd OOP vibration mode frequencies, being the higher results obtained by the panels with plaster. The variation of the 1st OOP vibration mode frequency of the panel with the different levels of axial load applied in the frame columns. It was observed that the axial load increased the panel OOP frequency and, consequently, the panel OOP stiffness around 38%. It was also carried out the study of the evolution of the OOP vibration mode frequencies along the period after construction, from which it was found that after the fourth day no significant variations were observed in both IP and OOP frequencies. Ambient vibration tests carried out in-situ, by testing infill panels from three different buildings. From the (preliminary) analysis of the studied cases results it can be concluded that the parameter that affects more the panels OOP frequency is the Linf /Hinf ratio and the openings. The reduction of the border constraints also reduces the OOP vibration frequency. From this testing campaign, is was also observed that double-leaf IM walls frequencies were about 2 times higher than the single leaf panels, but it was observed that the internal and external leaves obtained similar OOP vibration frequencies suggesting the presence of a connection between leaves.

4.5 Final Considerations

155

However, to complement and reinforce the global findings of the present study, additional tests should be performed in the future.

4.5 Final Considerations The characterization of the material, mechanical and dynamic properties of the existing infill panels are of full importance to understand the global seismic behaviour of the infill panels and to calibrate the numerical models. Since the infill panels’ design project is not mandatory, those properties are forgotten, which increases the difficulty to assess the seismic vulnerability of an existing RC building structure or to design a new one. In the literature, there is a reduced amount of studies concerning the characterization of the properties of the infill panels which was the major motivation for the development of the works herein presented in this Chapter. In this context, the first part of this chapter is an experimental campaign comprising the material and mechanical characterization wallets made with hollow clay horizontal brick units. Compression strength tests were carried out in individual masonry units with two different thicknesses (HCHB units 110 mm and 150 mm thick) from two different suppliers. The compressions strength of both units were obtained, from which it was found that the thinner ones reached the higher results. The mechanical characterization tests of masonry wallets comprised four different types of tests, namely: compression strength, diagonal tensile, flexural strength parallel and perpendicular to horizontal bed joints. The tests observed a fragile behaviour of the infill wallets due to the hollow clay horizontal bricks characteristics. The collapse of parts of the bricks, with the increment of the vertical compressive loads, increased the wallets’ OOP instability, resulting in fragile collapses. The plaster contributed to increment of the infill panels’ flexural strength. A table containing the material and mechanical properties of the wallets were provided. The second part of this study was related to the ambient vibration tests carried out in infill panels to obtain the 1st OOP vibration mode frequency. For this, laboratory and in-situ tests were carried out in infill panels with different characteristics (geometric dimensions, thickness, openings and boundary conditions). From the ambient vibration tests, it was observed that single-leaf panels’ first OOP frequency ranges from 15 to 36 Hz, regardless of the panel thickness. Depending on the typology and the opening area, the openings can reduce the OOP frequencies from 20–40%. The axial load applied on columns adjacent to the panel increased the panel OOP natural frequency and consequently the panel OOP stiffness, around 38%.

156

4 Experimental Characterization of the As-Built Masonry …

References 1. Furtado A, Rodrigues H, Arêde A (2018) Calibration of a simplified macro-model for infilled frames with openings. Adv Struct Eng 21(2) 2. Furtado A, Rodrigues H, Arêde A, Varum H (2018) Out-of-plane behavior of masonry infilled RC frames based on the experimental tests available: a systematic review. Constr Build Mater 168:831–848 3. Furtado A, Rodrigues H, Arêde A, Varum H, Grubiši´c M, Šipoš TK (2018) Prediction of the earthquake response of a three-storey infilled RC structure. Eng Struct 171:214–235 4. Furtado A, Costa C, Arêde A, Rodrigues H (2016) Geometric characterisation of Portuguese RC buildings with masonry infill walls. Europ J Environ Civil Eng, 1–16 5. Asteris PG, Cavaleri L, Di Trapani F, Tsaris AK (2017) Numerical modelling of out-of-plane response of infilled frames: state of the art and future challenges for the equivalent strut macromodels. Eng Struct 132:110–122 6. Trapani FD, Shing PB, Cavaleri L (2018) Macroelement model for in-plane and out-of-plane responses of masonry infills in frame structures. J Struct Eng 144(2):04017198 7. Furtado A, Rodrigues H, Arêde A, Varum H (2016) Experimental evaluation of out-of-plane capacity of masonry infill walls. Eng Struct 111:48–63 8. Furtado A, Rodrigues H, Arêde A, Varum H (2016) Simplified macro-model for infill masonry walls considering the out-of-plane behaviour. Earthq Eng Struct Dynam 45(4):507–524 9. Ferretti D, Michelini E, Rosati G (2015) Mechanical characterization of autoclaved aerated concrete masonry subjected to in-plane loading: experimental investigation and FE modeling. Constr Build Mater 98:353–365 10. Knox CL, Dizhur D, Ingham JM (2018) Experimental study on scale effects in clay brick masonry prisms and wall panels investigating compression and shear related properties. Constr Build Mater 163:706–713 11. CEN (2016) NP EN 771–1: Especificações para unidades de alvenaria Parte 1: Unidades cerâmicas (tijolos cerâmicos) 12. CEN (2004) EN 1015–11 Methods of test for mortar for masonry—Part 11: determination of flexural and compressive strength of hardened mortar 13. Singhal V, Rai DC (2014) Suitability of half-scale burnt clay bricks for shake table tests on masonry walls. J Mater Civ Eng 26(4):644–657 14. ASTM (2002) Standard test method for diagonal tension (shear) in masonry assemblages— ASTM E 519–02. American Society for Testing and Materials, Annual Book of ASTM Standard 15. RILEM (1994) RILEM TC 76-LUM. Diagonal tensile strength of small walls specimens. RILEM Publications SARL 16. CEN (1999) EN 1052–2: Methods of test for masonry. Determination of flexural strength 17. Pereira P,Pereira M, Ferreira J, Lourenço P (2012) Behavior of masonry infill panels in RC frames subjected to in plane and out of plane loads. 7th conference on on analytical models and new concepts in concrete and masonry structure Cracow, Poland 18. CEN (2004) Eurocode 2: design of concrete structures—Part 1–1: general rules and rules for buildings. CEN. Brussels: Eupean Comitee for Standardization. 19. Sinha BP, Pedreschi R (1983) Compressive strength and some elastic properties of brickwork. Int J Masonry Constr 3(1):19–27 20. Bartolomé S (1990) Colleción del Ingeniero Civil, (in Spanish). Libro No. 4 Colegio de Ingenierios del Peru 21. Hendry AW (1990) Structural masonry. Macmillan Education Ltd, London, England 22. Paulay T, Priestley M (1992). Seismic design of RC and masonry buildings. Wiley. ISBN: 0-471-54915-0 23. CEN (2005) Eurocode 6: Part 1–1—General Rules for buildings—Rules for reinforced and unreinforced masonry. European Committee for Standardisation, Brussels 24. Varum H, Furtado A, Rodrigues H, Dias-Oliveira J, Vila-Pouca N, Arêde A (2017) Seismic performance of the infill masonry walls and ambient vibration tests after the Ghorka 2015, Nepal earthquake. Bull Earthq Eng 15(3):1185–1212

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25. De Angelis A, Pecce MR (2018) Out-of-plane structural identification of a masonry infill wall inside beam-column RC frames. Eng Struct 173:546–558 26. NI (2016) National instruments (http://www.ni.com/) 27. SVS (2012) Structural vibration solution—ARTeMIS extractor pro. Release 5:3 28. Cunha Á, Caetano E (2006) Experimental modal analysis of civil engineering structures. Sound Vib 40(6):12–20 29. NI (2010) National instruments—labview signal express

Chapter 5

Experimental Evaluation of the Out-of-Plane Behaviour of As-Built Masonry Infill Walls

5.1 Introduction The infills OOP behaviour is a complex topic, since it depends on a high number of variables and uncertainties, which can result in high variability responses. There is a need to increase the number of tests to understand the effect of the different parameters in the global OOP behaviour. Over the literature, the number of OOP tests of infill masonry panels is reduced, which limits the understanding and findings concerning its non-linear behaviour [1]. Different loading approaches, different materials, different workmanships, different specimens’ detailing, and different scaling (especially the type of blocks) turns very difficult the extraction of solid conclusions. One of the major goals of this thesis is to characterize the OOP seismic vulnerability of infill panels representative of the built building stock in Portugal between 1970 and 2000 and the effect of different variables in their response. To accomplish this, an experimental campaign consisting of nine OOP quasi-static tests of as-built full-scale masonry infill walls was carried out. The objectives of this testing campaign were: (i) Development of a testing platform able to perform OOP tests of full-scale specimens; (ii) characterization the OOP behaviour and thus assessing the collapse vulnerability; and (iii) assessment of the effect of different variables such as OOP loading type (monotonic load or half-cyclic load), columns’ axial load, workmanship and previous damage due to prior IP test, panel width support, plastering and test setup. This experimental campaign is sub-divided into two different stages: Stage 1: six OOP tests carried out with nylon airbags, where the variables studied were OOP loading type, previous damage, columns’ axial load, infill support conditions and plaster. Stage 2: Three OOP tests carried out with pneumatic actuators, where the variables under study were the workmanship and the previous damage. The nine specimens were named Inf_# (where # is the number of test, namely # = 01, 02, …, 09 and 11). Throughout the present chapter, details regarding the specimens (geometry, materials and reinforcement detailing) will be first presented. © The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 A. F. Furtado, Seismic Vulnerability Assessment and Retrofitting Strategies for Masonry Infilled Frame Building, Springer Theses, https://doi.org/10.1007/978-3-031-20372-5_5

159

160

5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

Afterwards, the two stages of this testing campaign will be presented separately, in which the details concerning each test setup, instrumentation and loading protocol will be described. Furthermore, the results obtained from specimen test will be analysed in terms of force–displacement hysteretic curves, cracking pattern and observed damages. To sum up, a global comparison will be made between all the tests, to draw some conclusions regarding each variable under study. The results will be compared in terms of force–displacement envelopes, maximum OOP strength, stiffness degradation and cumulative energy dissipation.

5.2 Specimens Details The infill wall specimens’ geometric dimensions were defined as 4.20 × 2.30 m (length and width respectively, which are representative of those existing in the Portuguese building stock according to the study developed by [2], and shown in Fig. 5.1a and b. The columns’ and beams’ cross sections were considered as 0.30 × 0.30 m and 0.30 × 0.50 m, respectively. No seismic design was considered for the frame to represent the RC structures from the Portuguese National Code “Regulamento de Segurança e Ações (RSA)”, which is the first robust seismic design code in Portugal. Figure 5.1 shows the schematic layout of the RC frame geometry with the corresponding columns’ and beams’ dimensions and reinforcement detailing (Fig. 5.1c, d and e, respectively). All the infill panels have equal geometry with the abovementioned in-elevation dimensions, made of HCHB units with 150 mm thickness, as usually found in Portugal’s most common infill masonry. No reinforcement was used to connect the infill panel and the surrounding RC frame, and no gaps were adopted between the panel and the frame. All the panels, except specimen Inf_05 were built aligned with the external side of the RC beam. A summary of the experimental tests and corresponding main characteristics are shown in Table 5.1. From the testing campaign, the following considerations (by each specimen) can be added: • Inf_01: One-leaf panel (thickness: 150 mm); aligned with the external face of the supporting beam; OOP monotonic test, 300kN applied in the top of each column; total panel width supported in the bottom beam; no plaster; OOP load applied with airbags; • Inf_02: One-leaf panel (thickness: 150 mm); aligned with the external face of the supporting beam; OOP half-cyclic test; no additional vertical load applied; total panel width supported in the bottom beam; no plaster; OOP load applied with airbags; • Inf_03: Double-leaf panel, composed by one external leaf with thickness of 150 mm and an internal leaf of 110 mm; subjected to a previous IP test until reaching 0.5%; after the IP test the internal panel was removed and the external one was subjected to OOP half-cyclic test; no additional vertical load applied on

5.2 Specimens Details

161

(a)

(b)

(d)

(c) (e)

Fig. 5.1 Infilled RC frame specimen dimensions (units in meters): a general dimensions; b front view of the specimen; c RC frame reinforcement detailling; d column and e beam dimensions and reinforcement detailing

162

5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

Table 5.1 Comparative summary of the experimental tests performed Stage Specimen OOP loading type 1

2

*

•

• •

•

•

Previous Columns Infill panel Plaster in-plane axial support drift load conditions

Workmanship

Inf_01

Monotonic –

300

Full support

–

–

Inf_02

Half-cyclic –

–

Full support

–

–

Inf_03

Half-cyclic 0.5%

–

Full support

–

–

Inf_04

Half-cyclic –

270

Full support

–

–

Inf_05

Half-cyclic –

–

2/3 width support

–

–

Inf_06

Half-cyclic –

–

Full support

Yes/10 mm –

Inf_08

Half-cyclic –

–

Full support

Yes/10 mm –

Inf_09

Half-cyclic –

–

Full support

Yes/10 mm

Inf_11

Half-cyclic 0.3%

–

Full support

Yes/10 mm –

√

*

Last row of bricks applied by different team

the columns; total panel width supported in the bottom beam; no plaster; OOP load applied with airbags; Inf_04: One-leaf panel (thickness: 150 mm); aligned with the external face of the supporting beam; OOP half-cyclic test; vertical load of 270kN applied in the top of each RC column; total panel width supported in the bottom beam; no plaster; OOP load applied with airbags; Inf_05: One-leaf panel (thickness: 150 mm); aligned next to the external face of the supporting beam; half-cyclic test; no additional vertical load; 2/3 panel width supported in the bottom beam; No plaster; OOP load applied with airbags; Inf_06: One-leaf panel (thickness: 150 mm); aligned with the external face of the supporting beam; OOP half-cyclic test; no additional vertical load; total panel width supported in the bottom beam; 10 mm plaster; OOP load applied with airbags; Inf_08: One-leaf panel (thickness: 150 mm); aligned with the external face of the supporting beam; OOP half-cyclic test; no additional vertical load applied; total panel width supported in the bottom beam; 10 mm plaster; OOP load was applied with pneumatic actuators; Inf_09: One-leaf panel (thickness: 150 mm); aligned with the external face of the supporting beam; OOP half-cyclic test; no additional vertical load applied; total

5.3 Material Properties and Specimen Construction

163

panel width supported in the bottom beam; 10 mm plaster; different workmanship of the panel Inf_08; OOP load applied with pneumatic actuators; • Inf_11: One-leaf panel (thickness: 150 mm); aligned with the external face of the supporting beam; subjected to a prior IP cyclic test with a maximum drift equal to 0.3%; half-cyclic test; no additional vertical load applied; total panel width supported in the bottom beam; 10 mm plaster; OOP loading applied with pneumatic actuators.

5.3 Material Properties and Specimen Construction 5.3.1 Concrete and Reinforcement Steel Bars The concrete considered for the RC frame specimen construction consisted of regular C20/25-class concrete [3]. Table 5.2 summarizes the results of the corresponding compression tests for concrete strength and elastic modulus determination. Considering a correction factor of 0.8 to convert (approximately) the mean cubic compressive strength, the yield is fcm,cyl = 21.4 MPa. Then, assuming the approximate relation fck,cyl = fcm,cyl −8 MPa, which estimates the characteristic value of the compressive strength, fck,cyl = 13.4 MPa is obtained. For the RC frame specimen construction, three different bar diameters were used, from the same lot, namely 6, 10 and 16 mm. Three samples of each were taken from each diameter bar and tested according to [4]. The relevant results obtained are shown in Table 5.3, in terms of Young’s modulus (Es ), yield tensile strength (fsy ), ultimate tensile strength (fsu ) and ultimate tensile strain (εsu ), for each sample and the corresponding average value of each bar diameter. Table 5.2 Results from compression tests and elastic modulus determination tests on concrete specimens according to NP-EN206 2000 [3] Sample

Ultimate compressive strength fcu (MPa)

Average (MPa)

SD (MPa)

C.o.V (%)

Elastic modulus Ec,cil (GPa)

Average (GPa)

SD (MPa)

C.o.V (%)

1

26.3

26.8

2.1

8

24.5

24.7

0.8

3

2

28.8

24.3

3

26.8

24.1

4

28.2

26.0

5

27.7

25.3

6

22.9

24.0

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5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

Table 5.3 Results from tensile tests on steel bar specimens according to NP-EN10002-1 2006 [4] Diameter group

φ6 mm

φ10 mm

φ12 mm

Sample

Young’s modulus Es (GPa)

Yield strength fsy (MPa)

Ultimate strength fsu (MPa)

Ultimate strain εsu (‰)

1

208.7

450.3

570.5

16.2

2

205.3

441.5

619.9

17.7

3

198.5

440.3

589.1

15.3

Average

204.2

444.0

593.2

16.4

1

207.1

586.3

680.8

20.7

2

214.0

619.7

721.5

18.1

3

207.9

590.8

692.3

20.8 19.9

Average

209.7

598.9

698.2

1

203.7

492.8

616.6

26.9

2

212.4

510.5

632.5

24.6

3

212.2

479.8

595.9

27.2

Average

209.4

494.4

615.0

26.2

5.3.2 Mortar and Masonry Units As mentioned in this section, the mortar selected for the experimental studies herein presented in this thesis, is a traditional one, type M5 (“Ciarga”) which characteristics are representative of the mortar used in the construction practice of the Portuguese building stock in the 1970s and 1980s (Fig. 5.2a). Each bag, supplied by CIMPOR, brings specific recommendations regarding the water necessary for prepare the mortar for the construction of the panels (Fig. 5.2b). Flexural and compression strength mortar tests were carried out according to EN 196–2006 [4], six samples for each panel. The main results are summarized in Table 5.4, from which some variability of the mortar properties can be observed, namely the high compressive strength of the mortar used to build the panels Inf_01, Inf_03 and Inf_04, all of them with compressive strength higher than 10 MPa, differing to what is expected for a typical M5 class mortar. For this experimental campaign, two different suppliers provided the mortar: the wall panels Inf_01, Inf_02, Inf_03 and Inf_04 were built with the mortar provided by supplier 1; Inf_05 and Inf_06 were built with mortar provided by supplier 2. This can justify some of the variability found in the mortar properties and the high strength values obtained by the specimens built with the mortar provided by supplier 1. Concerning the results of the mortar used in stage 2, it can be noticed that again some slight variations were found, namely that the compressive strength of the mortar used for the construction of the panel Inf_11 was about 6.49 MPa, 1.24 times and 1.44 times higher than Inf_08 and Inf_09, respectively. The same was observed in the test results carried out in the mortar samples for the plaster.

5.3 Material Properties and Specimen Construction

165

(a)

(b)

Fig. 5.2 Mortar Type M5 (“Ciarga”) used to build the infill panels: a exterior view of the bag; and b general view of the mortar preparation

Table 5.4 Results from flexure and compressive strength tests on mortar specimens Stage

1

2

Wall panel

Inf_01

Mortar used for the construction of the infill panel

Mortar used for plaster

Flexure strength (MPa)

Compressive Curing time strength (MPa) (days)

Flexure strength (MPa)

Compressive strength (MPa)

5.65

16.55

–

–

82

Inf_02

2.11

5.66

114

–

–

Inf_03

4.27

13.40

118

–

–

Inf_04

4.51

11.91

47

–

–

Inf_05

1.79

4.45

49

–

–

Inf_06

2.47

7.75

50

1.82

7.10

Inf_08

1.90

5.24

43

1.81

4.01

Inf_09

1.72

4.49

55

1.92

6.96

Inf_11

2.22

6.49

50

2.32

8.20

Furthermore, during the construction of the wall panels it was visible by the technical staff the high sensibility of the mortar to small variations of water contents both in terms of visual appearance and workability. Even following the recommendations provided by the mortar technical sheet, it was found that the same quantity of water led to small differences between the mortars contained in each bag. A poor control of workmanship regarding the quantity of water used for each mixing could also justify

166

5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

some of the variations observed. Finally, it is included in Table 5.4 the curing time of the mortar specimens. All the mortar specimens were tested at the day of the infill panel OOP test, to ensure that the mortar properties were well characterized. HCHB brick units 150 mm thick were selected to build the infill panels. The material properties were presented in this section for the brick units from the group 1, supplied by PRECERAM. The panels were constructed after the full hardening of the RC frame. The thickness of both the adopted vertical and the horizontal bed joints was 10 mm. Full contact between the infill panel and the surrounding RC elements was achieved by filling the vertical and horizontal gaps between the infill and the frame with mortar. All the panels were built according to the traditional construction method, starting with applying the first layer of 10 mm mortar joint (interface panelbottom beam—Fig. 5.3a) and then placing the first brick next to the column. The proper filling of 10 mm vertical joint between the brick and the column was ensured. During the execution of each row of bricks, a correct alignment with the external face of the bottom beam was ensured (Fig. 5.3b). The infill panels were built on the same day, except the upper row of bricks (Fig. 5.3c) that was just made the next

(a)

(b)

(c)

(d)

Fig. 5.3 Construction methodology of the infill panels under test: a application of the first layer of mortar and the first brick next to the column; b execution of the first row of bricks; c construction of the panel penultimate row of bricks; and d execution of the last row of bricks

5.4 Stage 1—Out-of-Plane Tests Using Airbags

167

day to ensure proper hardening and settlement of the mortar bed-joints and avoid possible shrinkage cracks. Due to the geometric configuration of the frame, it was necessary to fill the upper row with half-height bricks and then the interface between the brick and the top beam was filled with mortar (Fig. 5.3d). Finally, mortar was again placed in each vertical head joint to ensure that all of them were correctly filled in both sides of the panel. The plaster was always applied a few days after the panel’s construction.

5.4 Stage 1—Out-of-Plane Tests Using Airbags 5.4.1 Test Setup The OOP tests carried out in Stage 1 consisted of the application of a uniformly distributed surface load through a system composed of seven nylon airbags, reacting against a self-equilibrated steel structure, as shown in Figs. 5.4 and 5.5. The application of a uniform OOP loading aims to mobilize the OOP response of the infill wall globally. In the literature, similar OOP load distribution adopted by other authors can be found [5, 6]. This reaction structure is composed of five vertical and four horizontal alignments of rigidly connected steel bars, in front of which a vertical wooden platform is placed to resist the airbag pressure and transfer it to the steel reacting grid elements. Thus, 12 steel threaded rods, crossing the RC elements in previously drilled holes, are used to equilibrate the reaction force resulting from the pressure applied by the airbags in the infill panel. The steel rods are strategically placed to evaluate the load distribution throughout the entire infilled RC frame resorting to load cells attached to each rod, which allowed continuous measurement of the forces transmitted to the reaction structure to which the rods were directly screwed. On the other extremity of each tensioned rod, appropriate nuts and steel plates are used to anchor the rod and to apply its reaction force to the concrete surface by uniformly distributed normal stresses, thus avoiding load concentration on the RC elements crossed by the rods. The axial load is applied in each column by means of a hydraulic jack inserted between a steel cap placed on the top of the column and an upper HEB steel shape, which, in turn, is connected to the foundation steel shape resorting to a pair of highstrength rods per column. Hinged connections were adopted between these rods and the top and foundation steel shapes; the axial load applied to the columns is continuously measured by load cells inserted between the jacks and the top of each column (Fig. 5.6). The pressure level inside the airbags was set by two pressure valves which are controlled according to the target and measured OOP displacement of the central point of the infill panel (the control node and variable) continuously acquired during the tests using a data acquisition and control system developed in the National Instruments LabVIEW software platform [7].

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5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

(a)

(b) 0

Strong floor

1

Foundation steel shape

2

High-strength rods (ø30mm) fixing the foundation

3

steel rod (ø20mm) connecting the RC frame to the

4

vertical high-strength rods (ø30mm) to apply axial

steel shape to the strong floor foundation steel shape load in columns 5

Steel cap

6

steel rods (ø20mm) connecting the RC frame and

7

distributing load plate

8

self-equilibrated reaction steel structure

9

counterweigh

the reaction structure

(c)

10

wood bars

11

hydraulic jack (for axial load application)

12

vertical wooden platform

13

Nylon airbags

14

Infill panel

15

RC column

16

Steel plate for rod force distribution

Fig. 5.4 Layout of the OOP test set-up: a front, b plan and c lateral view

5.4 Stage 1—Out-of-Plane Tests Using Airbags

(a)

169

(b)

Fig. 5.5 General view of the OOP experimental test set-up: a front view, b near view

Fig. 5.6 Test setup view: axial load application at the top of the RC frame columns

5.4.2 Instrumentation and Loading Protocol Half-cyclic OOP displacements were imposed with steadily increasing displacement levels, targeting the following nominal peak displacements: 2.5; 5; 7.5; 10; 15; 20;

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5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

25; 30; 35; 40; 45; 50; 50; 55; 60; 65 and 70 mm. Two half-cycles were repeated for each lateral deformation demand level at the control node. The central geometric point of the infill wall was selected as the control point since it was expected that is the region where the largest deformation of the panel will occur. The instrumentation assumed to monitor the tests was defined to measure the OOP displacements, panel rotation and resulted in a total number of twenty-one displacement transducers, thirteen of them DWT—Draw-Wire Displacement Transducer (blue circles—Fig. 5.7) and the remaining eight LVDT—Linear variable differential transformer (grey circles—Fig. 5.7). The DWT transducers were divided along three vertical and three horizontal alignments at the quarters of each vertical and horizontal dimension of the panel. The panel rotation was aimed to be measured through four pairs of LVDT placed at the corners of each infill panel. As previously stated, twelve load cells were used to monitor the loading forces in each steel rod connecting the RC frame to the reaction structure during the tests (Fig. 5.7). The axial load applied on the top of each column was kept constant and continuously monitored by two load cells inserted between the jacks and the top of each column.

5.4.3 Stage 1: Individual Test Results This sub-section presents the results of the stage 1 experimental campaign and discusses test by test. The OOP load (FOOP ) versus the OOP displacement in the centre of the panel (control node dOOP ) are shown, together with the damage evolution from the first observed cracks on the panel to the end of the test. The OOP drift was also computed related to the panel control node. Some considerations will be drawn regarding the first cracking (instant in which occur the first crack, maximum peak load and ultimate (residual) load.

5.4.3.1

Specimen Inf_01

The first cracking occurred for an OOP drift of 0.20% and an OOP total force equal to 49.12kN, corresponding to the beginning of detachment of the panel from the top beam (Fig. 5.8a). After that, the OOP strength increased until a maximum peak load equal to 75.93kN for an OOP drift of 1.82%. Vertical cracking plus the detachment from the top characterized the failure pattern of the panel at the end of the test, as shown in Fig. 8b. Other thinner cracks were also visible in the panel, as shown in Fig. 5.8b. In the back side of the panel, it was observed crushing of some bricks in the top and the bottom rows, due to the development of arching mechanism (Fig. 5.8c). With the increment of the OOP displacement, the cracking was more pronounced until the end of the test (due to the panel OOP instability) and due to the detachment of the panel from the top and bottom beams (Fig. 5.8d).

5.4 Stage 1—Out-of-Plane Tests Using Airbags

171

(a)

(b)

(c)

Fig. 5.7 OOP test instrumentation: a schematic layout; b Lateral view of the instrumentation; and c Front view of the instrumentation

The observed final cracking pattern is plotted in Fig. 5.8e, in which the red line represents the panel vertical cracking and the blue red lines represented the thinner cracks. The force–displacement curve is plotted in Fig. 5.8f.

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5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

(a)

(b)

(c)

(d) Out-of-plane drift at control node (%) 0.00 80

0.87

1.74

2.61

3.48

5.22

6.09

6.96

Inf_01 (Monotonic) Previous In-Plane drift: 0% Axial Load: 300 kN IM wall Fully suported No plaster Cracking point Peak Load Ultimate point

70

Out-of-plane force FOOP (kN)

4.35

60 50 40 30 20 10 0

End of test

0

10

20

30

40

50

60

70

80

Out-of-plane displacement at control node dOOP (mm)

(e)

Fig. 5.8 Inf_01: a First cracking; b Final cracking pattern; c Damages observed (Back view); d detail from the detachment of the panel from the bottom RC beam; e Force–displacement response

5.4.3.2

Specimen Inf_02

The first crack observed for an OOP drift of 0.11% and a corresponding total force equal to 51.09kN, is shown in Fig. 5.9a. After that, a slight detachment of the panel from the top beam triggered a trilinear crack, as shown in Fig. 5.9b. The panel

5.4 Stage 1—Out-of-Plane Tests Using Airbags

173

(a)

(b)

(c)

(d) Out-of-plane drift at control node (%) 0.00 80

0.87

1.74

2.61

3.48

5.22

6.09

6.96

Inf_02 (Half-cyclic) Previous In-Plane drift: 0% No axial load IM wall Fully suported No plaster Cracking point Peak Load Ultimate point

70

Out-of-plane force FOOP (kN)

4.35

60 50 40 30

End of test

20 10 0

0

10

20

30

40

50

60

70

80

Out-of-plane displacement at control node dOOP (mm)

(e)

Fig. 5.9 Inf_02: a First macro-cracking; b cracking pattern (Front view); c detail of the deformation in the centre of the panel; d damages observed (near view); e Force–displacement response

deformation was all concentrated in the centre of the panel, as visible in Fig. 5.9c. Figure 5.9d shows the back view of the panel, from which some bricks were crushing in the top of the panel due to the development of arching mechanism followed by the detachment from the top beam. The specimen Inf_02 achieved a maximum OOP strength of 6.94 kPa for an OOP drift of 1.44%. A residual strength equal to 37.80kN

174

5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

was reached for an OOP drift equal to 4.7%. The criterium to stop the test was to avoid collapse of the panel, which could damage the instrumentation located in the panel front. The cracking pattern of the specimen is plotted in Fig. 5.9b and the force– displacement response in Fig. 5.9e. Note, that the specimen Inf_02 is the reference specimen of Stage 1.

5.4.3.3

Specimen Inf_03

The specimen Inf_03, as mentioned before, was first subjected to a prior IP test until a maximum drift of 0.50%. During the IP test, a detachment of the panel from the top beam and from the top part of the columns was observed due to the in-plane frame distortion (Fig. 5.10a). When 0.40% IP drift was reached, corner crushing occurred in the left top corner of the panel, as shown in Fig. 5.10b. After the IP test, the panel Inf_03 was subjected to the OOP test and the behaviour was totally different than the previous ones, since for an OOP drift of 0.05% the panel detached from both the lateral columns and the top beam and started to behave as a rigid body (Fig. 5.11a and b). The panel cracking pattern is shown in Fig. 5.11c. A maximum peak load of 17.87kN occurred for a corresponding drift equal to 0.11%. Due to the OOP displacements profiles, the maximum ones occurred in the top of the wall; however, in Fig. 5.11d it is just plotted the maximum OOP displacement reached at the control node during the test. The ultimate stage of the test occurred for an OOP drift equal to 3.17% and a residual strength equal to 12.45kN (Fig. 5.11d). No strength degradation occurred during the post-peak stage, which can be explained by the strength provided by the friction of the top bricks and the top beam due to the global rotation of the panel.

(a)

(b)

Fig. 5.10 Inf_03: IP test a Damages observed (Front view); b detail of corner crushing

5.4 Stage 1—Out-of-Plane Tests Using Airbags

175

Fig. 5.11 Inf_03: a Observed damages (Front view); b detail of the panel detachment; c cracking pattern and d Force–displacement response

5.4.3.4

Specimen Inf_04

The behaviour of the specimen Inf_04 was different from the previous ones namely, the first cracking occurred for an OOP drift equal to 0.63% and the corresponding total force equal to 45.98kN, which was characterized by a vertical crack at the panel mid-span from top to bottom, as shown in Fig. 5.12a. The first macro-cracking also corresponded to the maximum peak load of the panel. After the peak load, there was a high decrease of OOP strength until an OOP force equal to 29kN which further reduced slightly until the ultimate load equal to 22.02kN and a corresponding drift equal to 3.91%. With the increment of the OOP displacements, the detachment of the panel from the top and bottom beams was visible, and the OOP force of the panel corresponded to the flexural mechanism of a panel constrained in 2 edges. The

176

5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

Fig. 5.12 Inf_04: a First cracking; b Observed damages; c cracking pattern and d Force–displacement response

damage observed during the test was the vertical crack located slightly to the left of the panel central alignment, from top to bottom. During the test, there was an increase of panel detachment from the top and bottom beams and the emergence of small, slightly inclined cracks near the panel corners (illustrated in Fig. 5.12b). The cracking pattern of the specimen is plotted in Fig. 5.12c and the force–displacement response in Fig. 5.12d.

5.4.3.5

Specimen Inf_05

In the specimen Inf_05, the first crack appeared for an OOP drift equal to 0.47% and a corresponding total force equal to 22.89kN; it was vertical and located in the middle of the panel (Fig. 5.13a). After that, the vertical crack extended to the top of the panel and from the middle of the panel to the bottom corners. The trilinear configuration, similar to the reference specimen Inf_02, is shown in Fig. 5.13b. The imposition of new displacements led that crack to spread towards the bottom corners of the masonry panel (Fig. 5.13c).

5.4 Stage 1—Out-of-Plane Tests Using Airbags

177

Fig. 5.13 Inf_05: a First cracking; b Lateral view of the damages observed; c cracking pattern; and d Force–displacement response

The partial support of the panel increased its OOP instability after the first crack appearance. In fact, for medium OOP displacements, panel instability was observed due to the reduced panel support width proved by the continuous sliding in the top of the panel and high residual displacements at the end of each half-cyclic. The maximum OOP strength of 27.82kN occurred for OOP drift of approximately 1.30%, while the ultimate OOP drift was about 1.83% without a pronounced decrease of the OOP strength, as proved by the ultimate strength equal to 22.99kN evidenced by the force–displacement response shown in Fig. 5.13d.

5.4.3.6

Specimen Inf_06

The panel Inf_06 failure was characterized by diagonal cracking from the central alignment of the panel towards the bottom corners. The cracking pattern is different from the ones observed in the previous specimens (which failure pattern was characterized by a trilinear and/or vertical cracking) because additional diagonal cracks were observed at the top of the wall. Besides the first cracking towards the left bottom corner (Fig. 5.14a), two diagonal cracks emerged from the middle of the panel to the

178

5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

Fig. 5.14 Inf_06: a First cracking; b Observed damages; c cracking pattern; and d Force–displacement response

right bottom corner (Fig. 5.14b). Slight detachment of the panel from the top beam was observed (Fig. 5.14b). The panel cracking occurred at 0.37% OOP drift, and after that the panel maximum strength was achieved by arching mechanism which highly contributed to the progressive increase of the OOP strength. After of the first cracking occurrence, the OOP total force suffered a reduction of 16%, after which it increased until the maximum force equal to 78.15kN associated with the ultimate OOP drift equal to 2.99%. The final cracking pattern of the specimen is plotted in Fig. 5.14c and the force–displacement response in Fig. 5.14d.

5.4.4 Stage 1: Global Results’ Comparison The present section aims to perform a comparative study between the OOP tests that were carried out in terms of cracking pattern and force–displacement curves from which it will be discussed some results parameters such as: OOP cracking strength and corresponding drift, OOP maximum strength and corresponding drift, ultimate OOP strength and corresponding drift. Finally, the results regarding the OOP displacement profiles, stiffness degradation and energy dissipation will also

5.4 Stage 1—Out-of-Plane Tests Using Airbags Table 5.5 Stage 1—Global results comparison: list of the tested infill panels and corresponding reference specimens

179

Wall panels

Reference specimen

Variants between tests

Inf_01

Inf_04

OOP loading type (monotonic/half cyclic) Mortar properties

Inf_02

OOP loading type (monotonic/half cyclic) Columns’ axial load Mortar properties

Inf_02

–

–

Inf_03

Inf_02

Previous damage Mortar properties

Inf_04

Inf_02

OOP loading type (monotonic/half cyclic) Columns’ axial load Mortar properties

Inf_05

Inf_02

Panel width support

Inf_06

Inf_02

Plaster

be presented and discussed. For a better understanding of the results discussion, Table 5.5 presents a summary list of the tested infill panels, the corresponding reference/comparative specimens, and the variants that could justify the differences among them.

5.4.4.1

Cracking Pattern

The specimen Inf_01 cracking pattern (Fig. 5.15a) was characterized by an approximate trilinear crack plus the detachment of the panel from the top and bottom RC beams. One diagonal crack developed from the central alignment of the wall towards the right bottom corner. However, the other expected diagonal in the left side of the panel did not occur. By comparing the Inf_01 cracking pattern with the one reached by the specimen with different loading protocol and characteristics (columns axial load, half-cyclic OOP load, full panel width support, no previous damage and no plaster) which is the specimen Inf_04 (Fig. 5.15d), one difference was observed, namely the pure vertical cracking achieved by the panel Inf_04. Both specimens, Inf_01 and Inf_04, cracking were followed by the detachment from the top and bottom frame’ beams. Some reasons could justify this difference, namely: (i) Inf_01 was subjected only to a monotonic OOP loading while the panel Inf_04 was subjected to a cyclic one; (ii) the mortar compressive and flexural strength used to build the panel Inf_04 are, respectively, 39% and 25% lower than the mortar used to build the specimen Inf_01; and (iii) the hardening time of the panel Inf_01 was 35 days larger than Inf_04. From these variables, it is not expected that the hardening time could affect the response of the specimens, since both were tested after the minimum hardening time

180

5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

Fig. 5.15 Stage 1—Comparative analysis: Cracking pattern a Inf_01; b Inf_02; c Inf_03; d Inf_04; e Inf_05 and f Inf_06

of 28 days. Thus, the mortar properties and the type of OOP loading (monotonic or cyclic) can justify the differences between panels Inf_01 and Inf_04 in terms of cracking patterns. Concerning to the differences observed between the panel Inf_01 and Inf_02 (Fig. 5.15b), these are related to the occurrence of detachment of the panel from the top and bottom frame’ beams (only in the case of specimen Inf_01) and the occurrence of a pure trilinear cracking pattern in the case of specimen Inf_02. In this case, this can be due to columns axial load effect which may have played an important role in the behaviour of panel Inf_01.

5.4 Stage 1—Out-of-Plane Tests Using Airbags

181

Regarding the comparison between the reference panel Inf_03 (subjected to halfcyclic OOP loading; no column axial load; with damage due to previous IP test (0.5% drift); full panel width support and no plaster) and the panel Inf_02 (subjected to halfcyclic OOP loading; no columns axial load; no previous damage; full panel width supported and no plaster), presented in Fig. 5.15c and b, it can be easily observed that the panel Inf_03 did not exhibit any type of cracking throughout the entire panel because the boundary conditions were severely affected during the IP test prior to the OOP test, in which led to the detachment of the wall from the envelope RC frame. Thus, when subjected to the OOP loadings, the panel already detached from the envelope frame and behaved as a rigid body. From this comparison, it seems clear that the previous damage caused by the IP demands was the key reason that led to the panel detachment, which increased the panel vulnerability to OOP loading demands. Regarding the comparison between the panels Inf_05 and Inf_02, a similar cracking pattern can be observed (Fig. 5.15e and b, respectively). Small differences between the two panels can be detected, namely a slight horizontal crack in the top of the panel Inf_05 plus a slight diagonal cracking in the right bottom side of the panel. Obviously, the major difference between them was the increased OOP instability of panel Inf_05 due to the reduced panel width support. In both tests, it was observed the arching mechanism development after the panel cracking. Finally, from the comparison between the specimens Inf_06 and Inf_02 (Fig. 5.15f and b), for which the main difference is the use of plaster, it can be detected the development of a larger number of cracks in the panel Inf_06. In the top of this panel small diagonal cracks were developed near the major vertical crack and also additional diagonal cracks emerged in the bottom part of the wall. The plaster apparently allowed the cracking distribution throughout the panel, which was not observed in the reference specimen Inf_02. To conclude, the OOP behaviour seems to depend on the mechanism that develops. On the other hand, the mechanism depends on a set of factors, such as: previous damage, panel support conditions, columns’ axial load, type of OOP load (monotonic/half-cyclic).

5.4.4.2

Force–Displacement Envelopes

From the analysis of the force–displacement envelope curves, plotted in Fig. 5.16, the following information was gathered for each tested panel: initial stiffness, OOP cracking strength and corresponding drift, OOP maximum strength and corresponding drift and OOP ultimate strength and corresponding drift, which will be presented and discussed in the next sub-sections.

Initial Stiffness By comparing the initial stiffness of all specimens, it can be noticed that the panel Inf_06 with plaster reached the highest initial stiffness and the specimen Inf_03

182

5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built … Out-of-plane drift at control node (%) 0.00 80

0.87

1.74

2.61

3.48

4.35

Inf_06

70

Out-of-plane force F OOP (kN)

5.22

6.09

6.96 Inf_01 Inf_02 Inf_03 Inf_04 Inf_05 Inf_06

Cracking point Peak Load Ultimate point

60 50 Inf_02

40 30 Inf_04

Inf_05

Inf_01

20 10

Inf_03

0 0

10

20

30

40

50

60

70

80

Out-of-plane displacement at control node dOOP (mm)

Fig. 5.16 Stage 1—Comparative analysis: Force displacement envelope curves

(with previous damage) the lowest one. The initial stiffness of the panel Inf_01 was 8% higher than that of Inf_02, which can be related to the columns axial load may have provided additional stiffness to the panel. The comparison between Inf_01 and Inf_04 found a 10% difference. As for the remaining test results, it was found that the initial stiffness of panel Inf_03 was 40% lower than the reference specimen Inf_02, which is justified by the previous damage. The comparison between the infill wall Inf_04 and Inf_02 shows that the first one was about 5% less stiff. Regarding the panel Inf_05 initial stiffness, it was reached a value 3% lower than the panel Inf_02, which suggests that the reduction of the panel width support did not affect significantly the panel initial stiffness. Finally, as previously mentioned, the specimen Inf_06 achieved the highest initial stiffness value, which was only 12% higher than Inf_02, basically due to the plaster contribution.

OOP Cracking Strength and Corresponding Drift Table 5.6 presents a summary of the OOP strength and drift values corresponding to the first observed crack. Additionally, it was included the ratio between each specimen result and the reference specimen Inf_02. The results found that the panel Inf_06 reached the highest OOP cracking strength and Inf_03 the lowest one with a value of 61.59kN and 16.88kN, respectively. From the comparison between the panel Inf_01 and Inf_02, a small difference around 4% is observed, however the OOP cracking drift was 1.81 times higher. The OOP cracking strength obtained by Inf_01 was 6% higher than that of the one reached by Inf_04. Again, the higher mortar compressive strength of the specimen Inf_01

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183

Table 5.6 Stage 1—Global results comparison: Cracking OOP drift and strength Specimen

OOP cracking strength FOOP,Crack (kN)

Ratio Inf_0#/Inf_02

OOP drift dOOP,crack (%)

Ratio Inf_0#/Inf_02

Inf_01

49.12

0.96

0.20

1.81

Inf_02

51.09

–

0.11

–

Inf_03

16.88

0.33

0.05

0.45

Inf_04

45.98

0.90

0.63

5.73

Inf_05

22.89

0.45

0.47

4.27

Inf_06

61.59

1.21

0.37

3.36

may have contributed to this result. Additionally, the OOP cracking drift of panel Inf_01 was about 70% lower than that of Inf_04. Looking for the result obtained by the panel Inf_03, the OOP cracking strength was 67% lower than that of Inf_02 and the OOP cracking drift occurred for a value 55% lower. It was clear that the previous damage affected both OOP cracking strength and drift results. The specimen Inf_04 reached an OOP cracking strength 10% lower than the panel Inf_02. In contrast with the result obtained by Inf_01, the columns axial load did not contribute to a higher OOP cracking strength. Nevertheless, the panel OOP cracking drift occurred in both panels, Inf_01 and Inf_04, for drift values higher than the one reached by Inf_02. Through the comparison between the Inf_05 and Inf_02 it can be noticed that for Inf_05 the OOP cracking strength was 55% lower, despite occurring for a drift value 4.27 times higher. The reduction of the panel width support reduced the OOP cracking strength capacity. Regarding the OOP cracking strength of the panel Inf_06, it was 21% higher than the value reached by Inf_02 and occurred for a drift 3.36 times higher. Globally, through the analysis of all the tests results it seems that the variables that affected more the panel OOP cracking strength are: (i) previous damage due to prior IP demands and (ii) reduction of the panel width support.

OOP Maximum Strength and Corresponding Drift Table 5.7 summarizes the results concerning the maximum OOP strength and corresponding drift reached by each specimen. Again, it was included the ratio between each specimen result and the reference specimen Inf_02. Furthermore, in Table 5.7 it was included the corresponding equivalent horizontal acceleration. The results found that the highest OOP maximum strength was reached by the infill panel Inf_06 and the lowest by Inf_03, corresponding to a value of 78.15kN and 17.87kN, respectively. Similarly, the same specimens again achieved the highest and lowest maximum OOP drift.

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5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

Table 5.7 Stage 1—Global results comparison: OOP maximum strength and corresponding drift Specimen

OOP maximum strength FOOP,max (kN)

Ratio Inf_0#/Inf_02

OOP drift dFmax (%)

Ratio Inf_0#/Inf_02

Inf_01

75.93

1.13

1.83

1.27

Inf_02

67.04

–

1.44

–

Inf_03

17.87

0.27

0.11

0.08

Inf_04

45.98

0.69

0.63

0.44

Inf_05

27.82

0.41

1.35

0.94

Inf_06

78.15

1.16

2.99

2.08

The specimen Inf_01 maximum OOP strength was 1.13 and 1.65 times higher than Inf_02 and Inf_04, respectively. Regarding the drift corresponding to the OOP maximum strength, for the panel Inf_01 it occurred for a drift value 1.27 times higher than Inf_02 and 2.90 times higher than Inf_04. In fact, the role played by the columns axial load is not very clear since the OOP maximum strength achieved by panel Inf_01 was higher than Inf_02 and, on the other hand, the value reached by Inf_04 was around 30% lower. Obviously, experimental variability can justify this difference, which means that a higher number of tests are needed to better understand the columns axial load effect in the maximum OOP strength of the panel. Concerning the effect of the previous damage, it can be concluded that the result of the specimen Inf_03 was 73% lower in terms of OOP maximum strength and 92% lower in terms of drift for which it occurred, comparing with Inf_02. Again, the strong impact of the previous damage in the panel OOP response is visible. Regarding the response of panel Inf_05, it can be noticed that the panel reached an OOP maximum strength 60% lower than the reference specimen and for a similar level of drift (around 5% of difference). It seems, as expected, that the reduction of the panel width support reduced the panel OOP capacity. Finally, as mentioned before, the plaster used in panel Inf_06 increased the OOP maximum strength capacity by about 16% and, besides that, it should be remarked that the drift for which it occurred was 2.08 times higher than the observed with the reference specimen. Concerning the computation and analysis of the equivalent acceleration, first it is needed to evaluate the mass of each infill panel. For that purpose a small infill panel was weighted, with geometric dimensions equal to 0.60 × 0.60 m, without plaster, which presented a total weight of 41 kg. By extrapolation, the corresponding infill panel mass per square meter (without plaster) is 114 kg/m2 . Assuming the total panel area equal to 9.66 m2 , the total panel mass is approximately equal to 1100 kg. Concerning the specimen Inf_06 mass (the unique panel with plaster), the plaster weight needed to be included in the total panel mass. The mortar self-weight was measured and is equal to 1758 kg/m3 . Considering that, the plaster thickness is equal to 1 cm, it results in a plaster total mass of 170 kg. Thus, the total infill panel mass (with plaster) is expected to be around 1270 kg. Additionally, based on the expected

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185

Table 5.8 Stage 1—Global results comparison: Equivalent horizontal acceleration Specimen

OOP maximum force (kN)

Total mass mpanel,total (kg)

66% Panel mass aequiv,tot (g)

mpanel,66% (kg)

aequiv,66% (g)

Inf_01

75.93

1100

0.68

726

1.03

Inf_02

67.04

1100

0.60

726

0.91

Inf_03

17.87

1100

0.16

726

0.24

Inf_04

45.98

1100

0.41

726

0.62

Inf_05

27.82

1100

0.25

726

0.38

Inf_06

78.15

1270

0.61

838

0.91

first-mode deformed shape, the mass participating in the first OOP vibration mode of the panel was also computed as 66% of the infill total mass, according to [8]. The equivalent horizontal acceleration was then obtained by dividing the maximum OOP load by the panel mass, and is described in Table 5.8 for all panels. It can be observed that the maximum value was obtained for the infill panel Inf_01 and the minimum one for Inf_03. The specimen Inf_06 equivalent horizontal acceleration is similar to the reference panel. Regarding the effect of the columns axial load, again two different results were obtained by Inf_01 and Inf_04, namely a 10% increase and a 34% decrease, respectively, compared to the result obtained by Inf_02. Similarly, to the previous results, the variables that apparently affect more the OOP maximum strength capacity is the previous damage and the reduction of the panel width support.

OOP Ultimate Strength and Corresponding Drift Table 5.9 shows the ultimate OOP strength and corresponding drift values obtained by each specimen. It should be reinforced that each test was stopped when it was visible the beginning of the panel instability to avoid the collapse, which means that the real capacity of the panel to resist to higher OOP displacement demands is not known. Thus, the conclusions regarding this parameter must have to be looked at in relative terms between the tested specimens, because the comparison between this parameter can only be reasonable and physically significant if the panels had reached the complete failure. From the results, the panel Inf_06 was the one that led to highest ultimate OOP ultimate strength (which corresponded also to the maximum OOP strength of all panels) and the panel Inf_03 the lowest one. However, regarding the corresponding OOP ultimate drifts, the maximum value was reached by the panel Inf_01 and the minimum one by the specimen Inf_05. The panel Inf_01 reached 51% and 12% lower OOP ultimate strength than Inf_02 and Inf_04, respectively. Regarding the panel Inf_03, the OOP ultimate strength was 67% lower than the one reached by Inf_02. Concerning the specimen Inf_05, the OOP ultimate strength was 42% lower than that of Inf_02. Finally, the specimen

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5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

Table 5.9 Stage 1—Global results comparison: OOP Ultimate strength and corresponding drift Specimen

OOP ultimate strength FOOP,ult (kN)

Ratio Inf_0#/Inf_02

OOP drift dOOP,ult (%)

Ratio Inf_0#/Inf_02

Inf_01

19.42

0.49

6.27

1.33

Inf_02

37.80

-

4.70

-

Inf_03

12.45

0.33

3.17

0.67

Inf_04

22.02

0.55

3.91

0.83

Inf_05

22.99

0.58

1.83

0.39

Inf_06

78.15

1.96

2.99

0.64

Inf_06 exceeded the Inf_02 ultimate OOP strength about 1.96 times. From these results, it appears that the plaster improved the OOP ultimate strength. On the other hand, the existence of previous panel damage, the reduction of the panel support width and the column axial load reduced the OOP ultimate strength. It has to be highlighted, again, the fact that this OOP ultimate strength and corresponding drift cannot be totally compared, and the effect of each variable cannot be extrapolated due to the reasons explained before. A possible comparison between the specimens is the conventional rupture, which corresponds to the reduction of the panel OOP maximum load about 20% (after peak stage). However, as observed in the force– displacement envelope curves, shown in Fig. 5.16, only three specimens reached this possible “conventional rupture”. This fact, reinforces again the difficulty inherent to the analysis of the ultimate stage of the tests.

OOP Displacement Profiles Aiming at comparing the OOP displacement profiles along vertical and horizontal alignments, their values were measured and compared for each test at the ultimate strength stage, Fult . For this analysis it were only considered three vertical alignments (L = 1/4Lpanel ; L = 1/2Lpanel ; L = 3/4Lpanel ) and only three horizontal alignments (H = 1/4Hpanel ; H = 1/2Hpanel ; H = 3/4Hpanel ). Due to the reduced number of displacement transducers, it is impossible to draw the complete vertical and horizontal profiles; however, with the results, it is possible to analyse and discuss the deformation of each tested panel. Regarding the vertical profiles (Fig. 5.17a–c), it can be observed that the larger OOP displacements occurred at the panel central alignment (L = 1/2Lpanel ), with particular focus for the infill panel Inf_02, which shows a linearly increasing OOP displacement along the panel height. The specimens Inf_03 and Inf_01, present a similar linear increasing of OOP displacements, however for lower values. In all the three vertical alignments, the panel Inf_05 reached lower OOP displacements, which could be because the panel was partially supported in the bottom beam.

1.6

1.4

1.4

Inf_01 Inf_02 Inf_03 Inf_04 Inf_05 Inf_06 Left alignment L=1/4Lpanel

1.2

1.0

0.8

1.6

Inf_01 Inf_02 Inf_03 Inf_04 Inf_05 Inf_06 Central alignment L=1/2Lpanel

1.2

1.0

0.8

0.6 30

40

50

60

70

80

10

20

30

40

50

60

OOP displacement (mm)

(a)

(b)

80

Inf_01 Inf_02 Inf_03 Inf_04 Inf_05 Inf_06 H=1/4Hpanel

70 60 50 40 30 20 10

70

1.4

1.6

1.8

2.0

2.2

2.4

Infill panel width (m)

(d)

2.6

2.8

3.0

0

10

20

30

40

50

60

70

80

OOP displacement (mm)

(c)

80

Inf_01 Inf_02 Inf_03 Inf_04 Inf_05 Inf_06 H=1/2Hpanel

70 60 50 40 30 20

80

H=3/4Hpanel

Inf_01 Inf_02 Inf_03 Inf_04 Inf_05 Inf_06

70 60 50 40 30 20 10

0 1.2

1.0

80

10

0

Inf_01 Inf_02 Inf_03 Inf_04 Inf_05 Inf_06 Right alignment L=3/4Lpanel

1.2

0.6 0

OOP displacement (mm)

OOP displacement (mm)

20

OOP displacement (mm)

10

1.4

0.8

0.6 0

OOP displacement (mm)

187

Infill panel heigth (m)

1.6

Infill panel heigth (m)

Infill panel heigth (m)

5.4 Stage 1—Out-of-Plane Tests Using Airbags

0 1.2

1.4

1.6

1.8

2.0

2.2

2.4

Infill panel width (m)

(e)

2.6

2.8

3.0

1.2

1.6

2.0

2.4

2.8

Infill panel width (m)

(f)

Fig. 5.17 Stage 1—Comparative analysis: OOP displacements vertical profiles at: a L = 1/4Lpanel ; b L = 1/2Lpanel and c L = 3/4Lpanel ; OOP displacement horizontal profiles at d H = 1/4Hpanel ; e H = 1/2Hpanel and f H = 3/4Hpanel

Regarding the horizontal OOP displacement profiles (Fig. 5.17d–f), it can be observed that the higher OOP displacements occurred at the H = 3/4Hpanel , and again the panel that reached higher and lower displacements were Inf_02 and Inf_05, respectively. From the analysis of the horizontal displacement profiles, it can be noticed that at H = 1/2Hpanel and H = 3/4Hpanel the deformation of the panel occurred at the central region of the panel, except the panel Inf_03 in which, as previously stated, it was observed a linearly increasing OOP displacement along the panel height. The OOP displacement contour level maps were plotted for each panel in Fig. 5.18, from which it can be observed the panel deformation and relate it with the observed

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5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

Fig. 5.18 Stage 1—Comparative analysis: OOP displacement contour level maps a OOP displacement transducers considered for the contour level maps b Inf_01; c Inf_02; d Inf_03; e Inf_04; f Inf_05 and g Inf_06

cracking pattern. The OOP displacement transducers used to plot the contour level maps of the panels Inf_01, Inf_02, Inf_04, Inf_05 and Inf_06 are indicated in Fig. 5.18a, which means that the panel part for which the contour map is plotted only along the horizontal range [1.05;3.05]m and the vertical range [0.575;1.725]m. The whole panel deformation was only possible to plot for specimen Inf_03, for which it was used additional instrumentation due to collaboration work with the INEGI research group regarding the use of Digital Image Correlation (DIC) technique for monitor the OOP displacements [9]. For that specimen, Fig. 5.18d fits to the whole panel geometric limits.

5.4 Stage 1—Out-of-Plane Tests Using Airbags

189

From the results, the panels can be divided into three groups: (a) more pronounced OOP deformation along the vertical alignment (Inf_01 and Inf_04); (b) largest OOP deformation in the top and central part of the panel (Inf_02, Inf_05 and Inf_06); and finally, (c) linear increasing of the OOP displacement along the panel height (Inf_03). Based on this, it seems that the columns axial load tends to contribute to the concentration of the panel deformation along a vertical alignment in the central part of the panel and that the previous damage due to the panel detachment tends to contribute to a rigid body behaviour.

Stiffness Degradation The stiffness degradation was evaluated by comparing the peak secant stiffness values obtained from the first cycle of each imposed peak displacement (Ks i+1 , Ks i+2 ,…, Ks i+n ) with the stiffness obtained at the 1st peak of displacement for all the panels (Fig. 5.19). The corresponding ratio was computed, leading to the relative stiffness plots shown in Fig. 5.20. From the results (Fig. 5.20), the following observations can be drawn: – As expected, due the nature of this type of elements, it is clear the trend of the stiffness degradation with the increase of the OOP displacements and corresponding loss of the internal integrity; – The lowest stiffness degradation for the same OOP displacement was obtained for panel Inf_01 (possibly because the panel was subjected to monotonic OOP loading combined with column axial load, which contributes to the increase of panel stiffness); on the other hand, the panel Inf_06 exhibited larger stiffness degradation for lower levels of OOP displacement; Out-of-plane drift at control node (%)

Fig. 5.19 Stage 1—Comparative analysis: methodology adopted to assess stiffness degradation

Ki+1 s

Ki+2 s

Out-of-plane force (kN)

Kis

Out-of-plane displacement at control node (mm)

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5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built … 1.0

Inf_01 Inf_02 Inf_03 Inf_04 Inf_05 Inf_06

0.9 0.8

Relative stiffness

0.7 0.6 0.5 0.4 0.3 0.2 0.1 0.0 0.00

0.87

1.74

2.61

3.48

4.35

5.22

6.09

6.96

OOP drift at control node (%) Fig. 5.20 Stage 1—Comparative analysis: global relative stiffness

– Despite the previous damage of the panel Inf_03, it is visible that it shows a similar stiffness degradation as that of Inf_04; – Although it is not fully supported, the panel Inf_05 evidenced a slower stiffness decrease compared to the panel Inf_02. Regarding the effect of the axial load on the top of the columns, no conclusions can be drawn, since the specimen Inf_01 was the one with lower stiffness degradation for the same OOP displacement, which was not observed for the specimen Inf_04 that exhibited behavior similar to the specimen Inf_03.

Cumulative Energy Dissipation For each infill panel, the energy dissipated in each individual loading half-cycle and the cumulative energy dissipation throughout the whole test history were calculated and plotted in Fig. 5.21. The cumulative dissipated energy was assessed for all the half-cyclic tests, considering the area of each loading half-cycle. The results show that the specimens Inf_02, Inf_03, Inf_04, Inf_05 and Inf_06 reached maximum cumulative energy dissipation values equal to 6kNm, 0.5kNm, 1.8kNm, 0.9kNm and 3.8kNm, respectively. It is not observed any visible trend (in all the tests) regarding the cycle for which the highest energy dissipation has occurred. It can be observed that the specimens Inf_03 and Inf_05 present similar energy dissipation capacity, which could mean that the reduction of the panel’ width support reduced the panel energy dissipation capacity due to the impossibility of the panel to reach higher

5.4 Stage 1—Out-of-Plane Tests Using Airbags

3

0.4

2 0.2 1 0.0 10

15

0 25

20

2 0.2 1 0.0 0

5

10

15

(b) 7 6

0.6

4 3

0.4

2 0.2 1 0.0 5

3

0.4

(a)

5

0

4

10

15

1.0

0 25

20

0.8

5 4 3

0.4

2 0.2 1 0.0 0

5

10

15

(c)

(d)

Dissipated energy for each cycle (kN.m)

6 5

0.6

4 3

0.4

2 0.2 1 0.0 5

10

15

20

0 25

7

Energy dissipated for each cycle Cumulative energy disspated Specimen Inf_06

0

6

0.6

Number of cycles

0.8

0 25

7

Energy dissipated for each cycle Cumulative energy disspated Specimen Inf_05

Number of cycles

1.0

20

20

Cumulative energy dissipated (kN.m)

0.8

5

Number of cycles

Energy dissipated for each cycle Cumulative energy disspated Specimen Inf_04

6

0.6

Number of cycles

Dissipated energy for each cycle (kN.m)

1.0

5

0.8

Cumulative energy dissipated (kN.m)

4

7

Energy dissipated for each cycle Cumulative energy disspated Specimen Inf_03

Cumulative energy dissipated (kN.m)

0.6

Dissipated energy for each cycle (kN.m)

5

0

Dissipated energy for each cycle (kN.m)

6

Cumulative energy dissipated (kN.m)

0.8

1.0

7

Energy dissipated for each cycle Cumulative energy disspated Specimen Inf_02

Cumulative energy dissipated (kN.m)

Dissipated energy for each cycle (kN.m)

1.0

191

0 25

Number of cycles

(e)

Fig. 5.21 Stage 1—Comparative analysis of individual half-cycle and cumulative energy dissipation a Inf_02; b Inf_03; c Inf_04; d Inf_05 and e Inf_06

levels of OOP displacement. Globally, the panel Inf_02, having the largest number of cycles, shows a higher energy dissipation in all of them, which contributed to a total energy dissipated around 11 times higher than the panel Inf_03 (specimen with lowest energy dissipation capacity). Figure 5.22 shows the cumulative energy dissipation evolution obtained for all the cyclic tested specimens. Again, it was found that the panel Inf_02 achieved the highest energy dissipation capacity for all the OOP displacement levels. On the

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5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

Cumulative energy dissipation (kN.m)

10

Inf_02 Inf_03 Inf_04 Inf_05 Inf_06

9 8 7 6

Inf_02

5 Inf_06

4 3 2

0 0.00

Inf_04

Inf_05

1

Inf_03

0.43

0.87

1.30

1.74

2.17

2.61

3.04

3.48

3.91

4.35

OOP drift at control node (%) Fig. 5.22 Stage 1—Comparative analysis: cumulative energy dissipation

other hand, it seems that the previous damage contributed significantly to reducing the capacity to dissipate energy. Concerning the specimen Inf_04, it can be concluded that it dissipated 70% less energy than the Inf_02. The comparison between the tests Inf_02 and Inf_05 shows that the first one dissipated significantly more energy than Inf_05 (around 6.7 times), even for all the cycles. The cumulative energy dissipation illustrated in Fig. 5.22, shows that the specimen Inf_02 dissipated more than 4 times higher energy than the remaining tests for the same OOP displacements. The panels Inf_04 and Inf_05 dissipated similar energy levels for the same OOP displacements. Globally, it can be observed that previous damage and the panel detachment from the surrounding frame reduces the infill panel capacity to dissipate energy and accommodate OOP forces. Besides that, it is observed that another variable which significantly reduces the panels’ energy dissipation capacity is the reduction of the panel support width.

5.5 Stage 2—Out-of-Plane Tests Using Pneumatic Actuators

193

5.5 Stage 2—Out-of-Plane Tests Using Pneumatic Actuators 5.5.1 Test Setup A novel test setup was developed, inspired by the setup described in this section, based in the application of a distributed OOP loading through several pneumatic actuators that mobilize the entire infill panel surface resorting to wood plates (one per actuator) placed between the actuators and the panel. One of the reasons behind this upgrade was to provide more stiffness and robustness to the reaction structure to support stronger infill walls. Additionally, this new testing platform can test infill panels with openings (with different configurations and dimensions) and was designed to allow performing combined IP-OOP quasi-static tests. This new test setup also allowed placing the instrumentation in the back of the reaction structure, which allows performing OOP tests until the panel collapse without damaging any equipment. In addition, the identification of panels’ cracking throughout the OOP tests becomes easier as also the quality of the pictures and films recording of each test improved significantly. In near future, it will be possible to apply different OOP loading patterns, alternatively to the traditional uniform shape. As mentioned before, this setup is based on the same concept of the previous one described in 6.4.1, namely the idea of a self-equilibrated system which balances the transmission of the OOP loadings to the reaction frame that, in turn is attached to the RC frame. Thus, this new test setup uses pneumatic actuators linked to four horizontal alignments performed by HEB140 steel shapes, which react against five vertical alignments made of HEB200 steel shapes. The horizontal alignments are coupled with hinged devices that allow lateral sliding. The steel structure is attached to the RC frame in twelve points (5 in the bottom beam, 5 in the top beam, and 2 in middle-height columns) with steel bars coupled with load cells to allow monitoring of the OOP loadings. Figure 5.23a and b show the schematic layout and the general view of the test setup. The pneumatic actuators performed the panel loading, the control is performed by monitoring their interior pressure and the imposed OOP displacements in a selected control point. General views of the testing setup are shown in Fig. 5.24.

5.5.2 Instrumentation and Loading Protocol Concerning the instrumentation adopted for the stage 2 tests, a few modifications were introduced namely: 25 displacement transducers were used to measure the OOP displacements of the panel along five horizontal and vertical alignments; the detachment of the panel from the surrounding corners was measured by LVDT’s placed in the transition infill plane-frame; 6 displacement transducers monitored the OOP displacements of columns and beams. Additionally, the vertical displacements of the

194

5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

(a)

(b)

Fig. 5.23 Layout of the OOP test set-up using pneumatic actuators: a front and b near view

5.5 Stage 2—Out-of-Plane Tests Using Pneumatic Actuators

195

(a)

(b)

(c)

(d)

Fig. 5.24 View of the OOP test set-up using pneumatic actuators: a overall view; b lateral view; c near view and d front view

top beam were measured by one LVDT placed at mid-span to monitor the evolution of possible arching mechanism (Fig. 5.25a). The support/reference structure for the monitoring system was placed behind the wall, as shown in Fig. 5.25b, which allow to carry out tests until panel collapse or until reaching the limit capacity of the pneumatics’ extension (100 mm). The loading protocol consisted of the application of half-cyclic OOP displacements (loading–unloading-reloading) that were imposed with steadily increasing displacement levels, targeting the following nominal peak displacements at the control node located in the centre of the panel: 0.5, 1, 1.5, 2, 2.5, 3.5; 5; 7.5; 10 mm; and then 5 by 5 mm up to a maximum OOP displacement of 120 mm. Two half-cycles were repeated for each lateral deformation demand level.

196

5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built … BV BT1 W1

W2

BT2 W6

h1 W3

W8

h2

W7 Legend

W5

W4

OOP masonry infill wall displacement

W10

W9

OOP RC frame displacement

h3

W13

W14

W16

h4

W17

W18

W20

h5

W22

C2

v5

v3

v1

W21 W19

W15

v4

W12

v2

W11

C1

W23

BB1

W24

W26 W25

BB2

(a)

(b)

Fig. 5.25 Stage 2: Instrumentation: a schematic front layout; and b general view

Beam Vertical displacement

5.5 Stage 2—Out-of-Plane Tests Using Pneumatic Actuators

197

5.5.3 Stage 2: Individual Test Results 5.5.3.1

Specimen Inf_08

The panel Inf_08, with plaster, subjected to half-cyclic OOP load, without columns axial load, without prior damage, was built by the workmanship A. Figure 5.26 shows the three different stages of the panel damages observed during the test, namely at first cracking (Fig. 5.26a) where it is visible the appearance of the central horizontal crack between the Hpanel /2 and HPanel /4. This first cracking occurred for an OOP drift equal to 0.11% and a corresponding load (FOOP,cr ) equal to 22.82kN. The second stage presented in Fig. 5.26b is related to the peak load one, where a trilinear cracking is visible with slight detachment of the plaster in the right and left sides of the panel and the beginning of detachment of the panel from the top beam. The peak load, equal to 44.22kN (1.93 times higher than FOOP,cr ) occurred for an OOP drift equal to 2.21%. The collapse of the panel was found to occur for an OOP drift equal to 3.20%, very close to the OOP drift at maximum peak load, and for a corresponding strength equal to 37.94kN (15% lower than the peak load), having occurred the detachment of the panel from the top and bottom beams, which led to a sudden loss of OOP stability and to the progression of the vertical crack until the bottom beam. The final damage visible at panel collapse is presented in Fig. 5.26c. Note that the panel did not collapse to the laboratory floor due to the existence of safety steel cables surrounding the frame to prevent that occurrence. The cracking pattern of this specimen, shown in Fig. 5.26d, is similarly to the panel Inf_06, a trilinear cracking with a slight horizontal cracks at Hpanel /4 and a two small vertical ones. It also indicated the length where the panel detachment from the envelope frame (top and bottom panel-frame interfaces) and some plaster detachment occurred. Figure 5.26a presents the OOP force–displacement response, where the 3 major stages are indicated (red circle—cracking point; blue circle—peak load; and green circle—panel collapse). The panel started with a continuous increase of strength with an initial secant stiffness equal to 21.68kN/mm. Afterwards, the first cracking was visible and, after that, the OOP displacement increased without any variation of the OOP load. The panel OOP strength increased until the peak load, after which there was a slight strength reduction about 15% until the panel collapse.

5.5.3.2

Specimen Inf_09

The panel Inf_09, with plaster, subjected to half-cyclic OOP load, without columns axial load, without prior damage, was built by the workmanship B. Figure 5.27 shows the different stages of the panel damages observed during the test, namely at first cracking (Fig. 5.26a) where again it is visible the appearance of the central horizontal crack between the Hpanel /2 and HPanel /4, which occurred for an OOP drift equal to 0.12% and a corresponding load equal to 44.25kN. Figure 5.27b shows the damages reached at the peak load stage, where a trilinear cracking is visible followed by

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5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

Fig. 5.26 Specimen Inf_08: a First cracking; b Cracking at Peak Load; c Final damage observed; d Cracking Pattern; and e force–displacement curve

the horizontal crack developed in the previous stage (extension of the first cracking towards the left and right columns). At this stage it was also visible the beginning of detachment of the panel from the top beam and some plaster detachment at the panel right side. The maximum peak load was 61.49kN, reached for an OOP drift equal to 2.54%.

5.5 Stage 2—Out-of-Plane Tests Using Pneumatic Actuators

199

Fig. 5.27 Specimen Inf_09: a First cracking; b Cracking at Peak Load; c last instant before the collapse; d beginning of the collapse mechanism; e panel collapse; f Cracking Pattern and g Force– displacement curve

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5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

The Fig. 5.27c–e present the last sequence of pictures just before the panel collapse, namely: immediately before the onset of the failure mechanism (Fig. 5.27c); the panel failure mechanism that triggered the collapse Fig. 5.27d, which is characterized by a trilinear division (cracking) of the panel and the detachment from the top and bottom frame beams, where it is also visible that the bottom part of the panel was divided into two parts, due to the progression of the vertical crack from the top to the bottom; finally, Fig. 5.27e presents the picture of the instant after the panel collapse with the detritus spread in the laboratory floor. The collapse occurred for an OOP displacement equal to 2.59% and an OOP load equal to 59.67kN. The cracking pattern is presented in Fig. 5.27f, characterized mainly by a trilinear cracking plus the slight cracks along the horizontal alignment at the 1/4 of the panel height. Furthermore, the detachment of the panel from the top beam was fundamental to trigger the collapse of the panel. Figure 5.27g shows the OOP force–displacement response plot, with and initial secant stiffness equal to 20.16kN/mm. The first cracking occurred as previously mentioned at 0.12% of drift and was followed by a continuous OOP strength increase until reaching the peak load equal to 61.49kN (40% higher than the OOP cracking strength) for an OOP displacement equal to 29.17 mm. Finally, a fragile collapse occurred without visible strength degradation as proved by the ultimate strength equal to 59.67kN (just 3% lower than the peak load).

5.5.3.3

Specimen Inf_11

Figure 5.28 shows the damages obtained after the IP test (with a maximum 0.3% drift) where it was observed the detachment of the panel from the top beam, the detachment of some plaster and a horizontal slide cracking in the left bottom of the panel. Details from the detachment of the panel from the top of the right column are shown in Fig. 5.28b and c. In the back side of the panel it was visible crushing of one brick in the top corner (Fig. 5.28d). No additional cracking was visible in the back side of the panel. Concerning the OOP test, the first cracking occurred at 0.67% drift and it was found vertical from the top of the panel down to mid-height, as shown in Fig. 5.29a, with a corresponding force equal to 23.5kN. The maximum peak load equal to 30.7kN was reached for OOP drift equal to 4.47% where a trilinear cracking pattern was visible with some slight cracks in the centre of the panel (Fig. 5.29b). The test was carried out until the maximum stroke of the pneumatic actuators, with an OOP displacement equal to 102.10 mm (drift equal to 8.87%) and a corresponding force of 13.3kN (Fig. 5.29c). The cracking pattern and the force–displacement response of the panel Inf_11 are shown in Fig. 5.29d and e, respectively.

5.5 Stage 2—Out-of-Plane Tests Using Pneumatic Actuators

201

Fig. 5.28 Specimen Inf_11: Damages after IP test a Cracking pattern; b detail of the plaster detachment; c detail of the panel-frame separation; and d near view with detail of the brick crushing

5.5.4 Stage 2: Global Results Comparison Similarly to 5.4.4, the present section aims to perform a comparative study concerning the tests carried out in stage 2. The comparison will be achieved by analysing the initial stiffness, cracking force and corresponding displacement, maximum peak load and corresponding displacement, ultimate strength and corresponding displacement, OOP displacement profiles, stiffness degradation and energy dissipation. The main variables herein discussed are the effect of the workmanship and the previous damage due the IP test. The major differences between the panel Inf_08 and the Inf_09 is that the panel Inf_09 upper horizontal bed joint (interface panel-top beam) was built by a different team than the remaining panels tested in this testing

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5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

Fig. 5.29 Specimen Inf_11: a First cracking; b Cracking at Peak Load; c Final observed damages; d Cracking pattern; and e force–displacement curve

campaign. For a better understanding of the results discussion, it is summarized in Table 5.10 the list of tested infill panels, corresponding reference specimens and the variables that could justify the differences among them.

5.5 Stage 2—Out-of-Plane Tests Using Pneumatic Actuators Table 5.10 Stage 2—Global results comparison: list of infill panels tested and corresponding reference specimens

203

Wall panels

Reference specimen

Variables between tests

Inf_08

–

–

Inf_09

Inf_08

Workmanship

Inf_11

Previous damage (IP test: 0.3%)

Fig. 5.30 Stage 2—Comparative analysis: cracking pattern a Inf_08; b Inf_09 and c Inf_11

5.5.4.1

Cracking Pattern

All the specimens cracking pattern, included in Fig. 5.30, were characterized by trilinear cracking with detachment of the panel from the top beam. In fact, both the workmanship and the previous damage did not introduce modifications in the panels’ cracking pattern. Some plaster detachment was observed in some of the panels.

5.5.4.2

Force–Displacement Envelopes

From the analysis of the force–displacement envelopes, plotted in Fig. 5.31, it is possible to observe different responses of the panels concerning different issues, such

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5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

Out-of-plane drift at control node (%) 0.00 0.87 1.74 2.61 3.48 4.35 5.22 6.09 6.96 7.83 8.70 9.57 8 80

Out-of-plane force FOOP (kN)

Inf_09

60

Panel collapse

50

Inf_08

7 6 5 4

40 Inf_09 Panel collapse

Inf_08

30

3

20

Inf_11

2

Inf_11

Out-of-plane pressure (kPa)

Inf_08 Inf_09 Inf_11 Cracking point Peak Load Ultimate Point

70

1

10 Limit of the pneumatic atuators

0 0

10

20

30

40

50

60

70

80

90

100

0 110

Out-of-plane displacement at control node d OOP (mm) Fig. 5.31 Stage 2—Comparative analysis: force displacement curves

as initial stiffness, cracking point (red circle), peak load (blue circle), ultimate point (green circle) and energy dissipation. From Fig. 5.31, it becomes evident that despite the strength reduction due to the previous damage, the specimen Inf_11 surprisingly reached higher OOP displacements than that of Inf_08, without collapse. On the other hand, both specimens Inf_08 and Inf_09 reached the collapse throughout the tests. In the following sub-sections the panels OOP response are presented and discussed in detail.

Initial Stiffness By comparing the initial stiffness, it is possible to observe that the panels Inf_08 and Inf_09 reached similar values, which revealed that the workmanship did not affect this aspect. Concerning the effect of the previous damage, the panel Inf_11 initial stiffness was 59% lower than the reference one. Clearly, the damage introduced by the IP test (detachment of the panel from the envelope frame plus some cracking and detachment of the plaster) affected as expect the initial stiffness of the panel as expected.

5.5 Stage 2—Out-of-Plane Tests Using Pneumatic Actuators

205

OOP Cracking Strength and Corresponding Drift By comparing the OOP cracking strength, the maximum value was reached by the specimen Inf_09, about 1.08 times higher than Inf_08, and the minimum one by the reference specimen Inf_08. Again, the previous damage did not affect the cracking strength since the panel Inf_11 reached a similar value than the reference one, but for a corresponding OOP drift 6.09 times higher. This could be justified by the existence of some cracking in the prior IP test. It is also important to remember that in the stage 1 for the specimen Inf_03 (subjected to previous 0.5% IP drift), it was not visible significant cracking apart of the detachment of the panel from the envelope frame (top beam and left and right columns). Table 5.11 presents the summary of the cracking OOP strength and corresponding drift obtained by all the specimens. Additionally, it is included the ratio between the values for specimens Inf_09 and Inf_11 and the reference one (Inf_08).

OOP Maximum Strength and Corresponding Drift Through the assessment of the maximum OOP strength achieved by the specimens (Table 5.12), it is possible to find some variations, namely the maximum one obtained by specimen Inf_09 with a value equal to 61.44kN, 1.39 times higher than for Inf_08. The specimen Inf_11 evidenced a reduction of the peak load to 30.72kN, about 30% lower than for Inf_08. From this, it can be reconfirmed that, as expected, the previous damage reduced the OOP strength capacity of the panel. Concerning to the OOP drift corresponding to the maximum strength, the result obtained by panel Inf_11 was 2.02 times higher than the reference one. Finally, the Table 5.11 Stage 2—comparative analysis: cracking oop drift and strength Specimen

OOP cracking strength FOOP,crack (kN)

Ratio Inf_#/Inf_08

OOP Drift dOOP,crack (%)

Ratio Inf_#/Inf_08

Inf_08

22.60

N/A

0.11

N/A

Inf_09

24.43

1.08

0.12

1.09

Inf_11

23.09

1.02

0.67

6.09

Table 5.12 Stage 2—comparative analysis: OOP maximum strength and corresponding drift Specimen

OOP maximum strength FOOP,max (kN)

Ratio Inf_#/Inf_08

OOP Drift dOOP,max (%)

Ratio Inf_0#/Inf_08

Inf_08

44.15

N/A

2.21

N/A

Inf_09

61.44

1.39

2.54

1.15

Inf_11

30.72

0.70

4.47

2.02

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5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

specimen Inf_09 achieved the maximum strength for an OOP drift equal to 2.54%, 15% higher than for Inf_08. From the analysis of the equivalent horizontal accelerations (Table 5.13), it can be observed that the maximum value was obtained by infill panel Inf_09 and the minimum one by Inf_11. The specimen Inf_09 reached an equivalent acceleration equal to 0.47 g (considering the total mass of the panel), 1.41 times higher than Inf_08 and 2.04 times higher than Inf_11. From that, it is clear that the previous damage strongly reduced the panel OOP strength capacity.

OOP Ultimate Strength and Corresponding Drift This new test setup allowed performing the OOP tests until the panel collapse and providing important information concerning the collapse mechanism, strength degradation, among others. From the results, the workmanship resulted in a variation of the ultimate strength of about 38%, with the higher result obtained by the Inf_09 with 59.70 kN. It was observed that both Inf_08 and Inf_09 collapsed without significant strength reduction, which can be related to the detachment of the panel from the top beam that may have triggered the failure mechanism responsible for the panel collapse. The panel Inf_09 collapsed for an OOP drift 20% lower than that one for which the reference specimen collapsed (Table 5.14).

OOP Displacement Profiles Similarly to what was performed in this section, i.e. the comparison of the vertical and horizontal profiles of OOP displacements, their values were measured, plotted and Table 5.13 Stage 2—global results comparison: equivalent horizontal acceleration Specimen

OOP maximum force (kN)

Total mass

66% Panel mass

mpanel,total (kg)

aequiv,tot (g)

mpanel,66% (kg)

aequiv,66% (g)

Inf_08

44.22

1270

0.34

838

0.52

Inf_09

61.49

1270

0.47

838

0.72

Inf_11

30.70

1270

0.24

838

0.36

Table 5.14 Stage 2—comparative analysis: OOP ultimate strength and corresponding drift Specimen

OOP ultimate strength FOOP,ult (kN)

Ratio Inf_0#/Inf_08

OOP ultimate drift dOOP,ult (%)

Ratio Inf_0#/Inf_08

Inf_08

37.96

N/A

3.20

N/A

Inf_09

59.70

1.57

2.59

0.81

Inf_11

13.33

0.35

8.87

2.77

5.5 Stage 2—Out-of-Plane Tests Using Pneumatic Actuators

207

Fig. 5.32 Stage 2—Comparative analysis: Vertical OOP displacements profiles at: a L = 1/4Lpanel ; b L = 1/2Lpanel and c L = 3/4Lpanel ; horizontal OOP displacement profiles at d H = 1/4Hpanel ; e H=1/2Hpanel and f H=3/4Hpanel

compared for each test at the ultimate strength stage, FOOP,ult . For this analysis it were only considered three vertical alignments (L = 1/4Lpanel ; L = 1/2Lpanel ; L = 3/4Lpanel ) and three horizontal alignments (H = 1/4Hpanel ; H = 1/2Hpanel ; H=3/4Hpanel ). Regarding the vertical profiles of displacements, plotted in Fig. 5.32, it is possible to observe that specimens Inf_08 and Inf_09 behaved similarly, with the maximum OOP displacements located in the panel mid-height and the lower values achieved by the panel Inf_09. Again, the OOP displacement values reached by panel Inf_11 were much higher than the remaining ones at the panel mid-height and top. Even the profile is different, since the higher OOP displacements occurred in the top of the panel, which is related to the panel detachment from the envelope frame caused by the previous damage.

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5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

Fig. 5.33 Stage 2—Comparative analysis: OOP displacement contour level maps a Inf_08; b Inf_09; and c Inf_11

Finally, concerning the horizontal profiles, all the panels achieved the higher OOP displacements in the panel mid length. The values reached by Inf_08 and Inf_09 were again similar and also much lower than those of Inf_11. The OOP displacement contour level maps are plotted for each panel in Fig. 5.33, from which the deformation of the panel can be observed and relate with the observed cracking pattern.

Stiffness Degradation The stiffness degradation was analysed for each specimen and compared with the remaining panels and is shown in Fig. 5.34. The panel Inf_08 achieved the largest stiffness decrease until an OOP displacement equal to 7 mm, where the stiffness reduction was higher than 80%. For OOP displacements larger than 7 mm, the degradation was softer. The panel Inf_09 presented a similar behaviour concerning the

5.5 Stage 2—Out-of-Plane Tests Using Pneumatic Actuators

209

Out-of-plane drift at control node d OOP (%) 0.00 0.87 1.74 2.61 3.48 4.35 5.22 6.09 6.96 7.83 8.70 9.57

Inf_08 Inf_09 Inf_11

1.0

Relative stiffness

0.8

0.6

0.4

0.2

0.0 0

10

20

30

40

50

60

70

80

90

100 110

Out-of-plane displacement at control node d OOP (mm) Fig. 5.34 Comparative analysis: global relative stiffness

behaviour before and after 7 mm, which curiously is the displacement where the first cracking occurred. However, when compared with the Inf_08 curve, the reduction is slightly lower. Finally, since the panel Inf_11 started the OOP test with prior damage, which means lower initial stiffness, it has presented slighter reduction of stiffness throughout the test when compared with the remaining ones. In the case of specimen Inf_11 three different stages can be identified, namely: higher stiffness degradation from the beginning of the test until an OOP displacement equal to 10 mm; an intermediate stage between 10 and 22 mm and a third stage for OOP displacements larger than 22 mm.

Cumulative Energy Dissipation The energy dissipated in each individual loading half-cycle and the cumulative energy dissipation throughout the whole test history were calculated and plotted in Fig. 5.35. The specimen Inf_11 reached the maximum cumulative energy dissipation, equal to 3.95kNm, 2.07 times higher than that of Inf_08 which dissipated 1.91kNm. The panel Inf_08 reached a cumulative energy dissipation 25% higher than the specimen Inf_09 did. Obviously, the fact that the panel Inf_11 reached higher OOP displacements allowed to dissipate high levels of energy throughout the test. To conclude, Fig. 5.36 shows the cumulative energy dissipation of all the specimens, from which it can be observed that for the same OOP displacement, the panel

5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

0.40

4.0

0.35

3.5

0.30

3.0

0.25

2.5

0.20

2.0

0.15

1.5

0.10

1.0

0.05

0.5

0.00 0

10

20

30

40

50

5.0

0.50

Dissipated energy for each cycle (kN.m)

Energy dissipated for each cycle Cumulative energy disspated 4.5 Specimen Inf_08

Cumulative energy dissipated (kN.m)

Dissipated energy for each cycle (kN.m)

0.45

0.0 60

Energy dissipated for each cycle Cumulative energy disspated 4.5 Specimen Inf_09

0.45

4.0

0.40 0.35

3.5

0.30

3.0

0.25

2.5

0.20

2.0

0.15

1.5

0.10

1.0 0.5

0.05 0.00 0

10

20

30

Number of cycles

Number of cycles

(a)

(b)

Dissipated energy for each cycle (kN.m)

0.50

4

0.40 0.35

3

0.30 0.25 0.20

2

0.15 1

0.10 0.05 0.00 0

10

20

30

50

0.0 60

5

Energy dissipated for each cycle Cumulative energy disspated Specimen Inf_11

0.45

40

Cumulative energy dissipated (kN.m)

5.0

0.50

40

50

Cumulative energy dissipated (kN.m)

210

0 60

Number of cycles

(c)

Fig. 5.35 Stage 2—Comparative analysis: individual half-cycle and cumulative energy dissipation a Inf_08; b Inf_09 and c Inf_11

Inf_09 dissipated higher levels of energy (1.55 times when compared with Inf_08 and 2.30 times than that of Inf_11) and the specimen Inf_11 dissipated the lowest values. This again, can be justified by the previous damage of the panel Inf_11. The responses of the Inf_08 and Inf_09 are significantly different.

5.6 Overview of the Summary of Global Results: Critical Analysis Concluded the presentation of the experimental campaign, which comprised two different stages of quasi-static OOP tests, it is important to discuss the global results and highlight the major findings. Table 5.15 presents the summary of the cracking, maximum and ultimate OOP strength, from which is possible to observe that the highest OOP cracking strength was obtained by Inf_06 and the lowest by Inf_03, with 61.59kN and 19.84kN, respectively. Specimen Inf_06 achieved the highest maximum peak load (half-cyclic OOP load; without previous IP damage, without columns axial load; with plaster) and the lowest by Inf_03 (half-cyclic OOP load; without previous IP damage, with columns axial

5.6 Overview of the Summary of Global Results: Critical Analysis OOP drift at control node (%) 0.00 5.0

Cumulative energy dissipation (kN.m)

Fig. 5.36 Stage 2—Comparative analysis: cumulative energy dissipation

211

1.74

3.48

5.22

6.96

8.70

10.43

40

60

80

100

120

Inf_08 Inf_09 Inf_11

4.5 4.0 3.5 3.0 2.5 2.0 1.5 1.0 0.5 0.0 0

20

OOP displacement at control node (mm)

Table 5.15 Summary of the global results cracking, maximum and ultimate OOP strength Stage Specimens FOOP,crack (kN) FOOP,max (kN) FOOP,ult (kN) Type Previous Plaster of IP OOP damage load Stage Inf_01 1 Inf_02

49.12

75.93

19.42

M

No

No

51.09

67.04

37.80

HC

No

No

Inf_03

16.88

17.87

11.69

HC

Yes (0.5%)

No

Inf_04

45.98

45.98

22.02

HC

No

No

Inf_05

22.89

27.82

22.99

HC

No

No

Inf_06

61.59

78.15

78.15

HC

No

Yes

Stage Inf_08 2 Inf_09

22.60

44.15

37.96

HC

No

Yes

24.43

61.44

59.70

HC

No

Yes

Inf_11

23.09

30.72

13.33

HC

Yes (0.3%)

Yes

M—Monotonic OOP load HC—Half-cyclic OOP load

load; without plaster), with 78.15kN and 17.87kN, respectively. Last but not least, the ultimate OOP strength was also included in Table 5.15, however this value cannot be compared among the specimens since the criteria adopted was not the same for all the panels. This parameter will be only possible to assess when all the panels are tested until their collapse. Table 5.16 presents the summary of the OOP drift corresponding to the cracking, maximum and ultimate load. Starting from the OOP drift corresponding to the

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5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

Table 5.16 Summary of the global results – OOP drift corresponding to the cracking, maximum and ultimate OOP strength Stage Stage 1

Stage 2

Range

Specimens

dOOP,crack (%)

dOOP,max (%)

dOOP,ult (%)

Inf_01

0.20

1.83

6.27

Inf_02

0.11

1.44

4.70

Inf_03

0.05

0.11

3.17

Inf_04

0.63

0.63

3.91

Inf_05

0.47

1.35

1.83

Inf_06

0.37

2.99

2.99

Inf_08

0.11

2.21

3.20

Inf_09

0.12

2.54

2.59

Inf_11

0.67

4.47

8.87

0.11–0.63%

0.63–2.99%

2.59–4.70%

cracking strength, the highest value and lowest value were obtained with specimen Inf_11 and Inf_03, respectively, 0.67% and 0.05%. Those specimens were the ones with previous damage due to IP tests. Inf_06 reached the highest OOP drift corresponding to the maximum peak load with 2.99% while the lowest was obtained from Inf_03 with 0.11%. From the comparison between the specimens Inf_06 (half-cyclic OOP load; without columns axial load; no previous damage; with plaster; OOP load with airbags) and Inf_08 (half-cyclic OOP load; without columns axial load; no previous damage; with plaster; OOP load with pneumatic actuators), it is possible to observe different responses, namely the cracking, maximum and ultimate strength reached by the specimen Inf_06 was 64%, 44% and 52% lower. The corresponding drift at cracking and maximum peak load reached by specimen Inf_06 were also 71% and 27% lower. A large variation of results were found concerning this response parameter which evidences the large variability and difficulty to characterize the infill panels’ response due to their high non-linearity. Lastly, the OOP drift corresponding to the ultimate strength of the specimens are also summarized in Table 5.16. Figure 5.37 shows the comparison among the tested specimens, where it is plotted in red the values corresponding to the appearance of the first cracking, in green to the maximum peak load and in blue to the ultimate stage of the tests. From the testing campaign, six different variables were studied, namely: • • • • • •

Effect of column axial load; Effect of reduction of the panel’ support width; Effect of previous damage; Effect of plaster; Effect of workmanship; Effect of test setup.

5.6 Overview of the Summary of Global Results: Critical Analysis 10

100 FOOP,crack

90 80

FOOP,max

9

FOOP,ult

8

dOOP,crack dOOP,max dOOP,ult

7

70

OOP drift dOOP (%)

OOP strength FOOP (kN)

213

60 50 40 30

6 5 4 3

20

2

10

1 0

0 Inf_01 Inf_02 Inf_03 Inf_04 Inf_05 Inf_06 Inf_08 Inf_09 Inf_11

Inf_01 Inf_02 Inf_03 Inf_04 Inf_05 Inf_06 Inf_08 Inf_09 Inf_11

Specimens

Specimens

(a)

(b)

Fig. 5.37 Summary of the global results—a OOP strength (FOOP ); and b OOP drift (dOOP )

The next sub-sections will describe the major findings concerning the effect of each variable the panel OOP behaviour of the infill panels.

5.6.1 Effect of Columns Axial Load The axial load applied in the top of the adjacent columns caused the modification of the cracking pattern (Fig. 5.38a and b), since the panel without axial load (Inf_02) evidenced a trilinear cracking with slight detachment of the panel from the top beam for high levels of OOP displacements. Concerning the panel Inf_04, a pure vertical crack at the panel central alignment was visible because of the detachment of the panel from the top and bottom beams. The effects of the force–displacement curves (Fig. 5.38c) are noticeable since the axial load caused a reduction of the panel strength capacity. Starting from the OOP strength corresponding to the first cracking, the axial load reduced about 39% the OOP strength, however it occurred for an OOP drift 1.27 times higher. The maximum peak load reduced about 32%, and again it occurred for an earlier OOP displacement. The post-peak stage, the specimens response were different since for the panel Inf_02 it is visible a pronounced strength degradation and while for the panel Inf_04 a sudden strength degradation occurred (at the moment when the vertical crack formed) after which a slight reduction of the panel OOP strength was noticed. Regarding the cumulative energy dissipation, the axial load reduced the capacity of the panel to dissipate energy for the same OOP displacement level. The total cumulative energy dissipation of the panel Inf_04 was 75% lower than the panel without axial load in the adjacent columns.

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5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

Fig. 5.38 Effect of the gravity load: a Inf_02 cracking pattern; b Inf_04 cracking pattern; c Force– displacement curves; and d cumulative energy dissipation

5.6.2 Effect of Reduction of the Panel Support Width This variable was studied in this testing campaign, and a panel with 2/3 width supported was tested and compared with the reference specimen Inf_02. Concerning the cracking pattern, it is visible that it was not affected by the reduction of the panel support width, as shown in Fig. 5.39a and b. Besides the appearance of additional minor cracks, no significant cracks was reported during the test. Regarding the force–displacement curves (Fig. 5.39c), the effect of reducing the panel support width becomes more evident since it reduced the strength and deformation capacity. The first cracking appeared for a similar OOP drift, but the corresponding OOP strength was 62% lower than the reference specimen. The maximum peak load reached by the specimen Inf_05 was 59% lower than the panel Inf_02 and, again, it occurred for similar OOP drift demands. As expected, the reduction of the panel support width reduced the deformation capacity of the panel which is the reason for having stopped the OOP test for low OOP displacement levels. It was also visible that it did not occur any significant strength degradation in the post-peak stage, as opposite of what occurred in the Inf_02 test. Finally, the reduction of the panel support width reduced

5.6 Overview of the Summary of Global Results: Critical Analysis

215

Fig. 5.39 Effect of the reduction of panel support width: a Inf_02 cracking pattern; b Inf_05 cracking pattern; c Force–displacement curves; and d cumulative energy dissipation

the capacity of the panel to dissipate energy for the same OOP displacement levels. Figure 5.39d shows that for the ultimate OOP displacement of Inf_05, this panel dissipated less 54% energy than the reference specimen did.

5.6.3 Effect of Previous Damage The study of the effect of the previous damage in the OOP capacity was one of the focus in this testing campaign. Due to the high costs associated (damages imposed in the RC frame) it becomes very difficult to carry out full-scale tests of panels previously damaged. In this testing campaign, two levels of previous IP drift were imposed: 0.3% and 0.5%, respectively Inf_11 and Inf_03 (note that the first one was built with plaster and the last without). Over the literature, it is quite acceptable that the first appearance of cracking due to IP loadings ranges between 0.15–0.25% and the cracking phase can be extended until 0.50%. After that, it is considered too extensive level of damage. Ideally, it would be important to carry out OOP tests in panels subjected to several different prior IP tests with incremental drifts. However,

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5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

this was not possible and it was decided to carry out OOP tests in one panel with low prior IP drift equal to 0.30% and in other with medium prior IP drift equal to 0.50%. From the result of the specimen tested with low prior IP drift (Inf_011), a trilinear cracking pattern was observed similar to the one of the reference specimen (Inf_08), as shown in Fig. 5.40a and b. However, in the specimen Inf_08 it was found a large number of slighter cracks over the panel, namely a horizontal one in the bottom part of the panel. The detachment of the panel Inf_11 from the top beam occurred for earlier OOP displacement demands, which is justified by the impact of the previous IP damage that results in the detachment of the panel from the top beam. In the case of the reference specimen, the detachment of the panel from the top and bottom beams occurred only few instants before the panel collapse. The previous damage reduced the maximum OOP strength about 31% and this occurred for an OOP drift equal to 4.47%, which is 2.02 times higher than the one obtained in the reference specimen. Surprisingly, the panel Inf_08 collapsed for more reduced OOP drift demands and on the other hand, it was not possible to test the panel Inf_11 until its collapse due to the limitation of the pneumatic actuators. This result can be justified by experimental variability and certain phenomena that allowed developing arching mechanism and higher deformation capacity. To finish,

Fig. 5.40 Effect of the previous damage: a Inf_08 cracking pattern; b Inf_11 cracking pattern; c Force–displacement curves; and d cumulative energy dissipation

5.6 Overview of the Summary of Global Results: Critical Analysis

217

by analyzing the cumulative energy dissipation it can be noticed that the panel Inf_11 dissipated lower energy than the panel Inf_08 for the same OOP displacement levels (Fig. 5.40d). For example, at the collapse moment of the panel Inf_08, the energy dissipated was 1.79 times higher than the one reached by Inf_11. Regarding the specimen subjected to previous medium damage (Inf_03), a failure mode was observed completely different from the reference one. The specimen behaved as a rigid body and detached from the envelope frame without any visible cracking, showing a fragile behaviour. On the other hand, the reference panel evidenced a trilinear cracking pattern combined with slight detachment between the panel and the top beam. The cracking pattern of each specimen is shown in Fig. 5.41a and b. Concerning the force–displacement responses (Fig. 5.41c), the comparison highlights the impact of the previous damage in the OOP response of the panel, namely by evidencing a reduction of the maximum peak load about 74%, without any degradation of strength during the post-peak stage. Obviously, the strength corresponding to the appearance of the first cracking was also 67% lower and for an early OOP displacement. Finally, the evolution of the cumulative energy dissipation, shown in Fig. 5.41d, presents a reduction of about 91%, which reinforce the idea that with the increase of the prior damage the panel OOP vulnerability increases.

Fig. 5.41 Effect of the previous damage: a Inf_02 cracking pattern; b Inf_03 cracking pattern; c Force–displacement curves; and d cumulative energy dissipation

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5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

5.6.4 Effect of Plaster The effect of the plaster was evaluated with the panel Inf_06, considering 10 mm plaster and compared with the reference one Inf_02 without plaster. From the cracking pattern, shown in Fig. 5.42a and b, it is possible to observe that the plaster allowed to distribute the cracking over the panel apart from the trilinear pattern that was possible to identify. From the force–displacement curves, shown in Fig. 5.42c, the strength corresponding to the first cracking is similar but it occurred for an OOP drift slightly higher in the specimen Inf_06. The plaster contributed to the increase of the OOP maximum peak strength about 1.17 times. No conclusions can be drawn regarding the ultimate stage, since the stopping criteria was not the same for both panels. The criteria was based on damage visualization and assessment to independently prevent the panel’s OOP collapse in each case. About the evolution of the cumulative energy dissipation, shown in Fig. 5.42d, it was observed that the plaster did not contributed to its increase. On the contrary, the panel Inf_06 result was 70% lower.

Fig. 5.42 Effect of the plaster: a Inf_02 cracking pattern; b Inf_06 cracking pattern; c Force– displacement curves; and d cumulative energy dissipation

5.6 Overview of the Summary of Global Results: Critical Analysis

219

5.6.5 Effect of Workmanship One variable under study was also the workmanship, since it is one of the issues with higher difficulty to control and is expected to be one source of huge uncertainty with impact in the panels’ behaviour. The variability associated with the workmanship in the construction of one infill panel is important and to assess that variable, two panels were constructed by two different teams (note that only the row of bricks and bed joint interface of the panel Inf_09 was built by a different team; the remaining rows of that panel and the panel Inf_08 was built by the same team). During the test, a trilinear cracking pattern was obtained in both panels (Fig. 5.43a and b), combined with some plaster detachment in both sides of the panel and some minor cracking (mainly horizontal cracks). As for the force–displacement curves, shown in Fig. 5.43c, it can be observed that the workmanship led to very significant variations of the panel response, both in terms of the moment when it occurred the first cracking, the maximum peak load and the panel collapse. Variations 20%-30% were found and the collapse occurred for an OOP displacement 20% lower. Finally, the variation concerning the cumulative energy dissipation for the same OOP displacement was around 35%, as shown in Fig. 5.43d.

Fig. 5.43 Effect of the workmanship: a Inf_08 cracking pattern; b Inf_09 cracking pattern; c Force–displacement curves; and d cumulative energy dissipation

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5 Experimental Evaluation of the Out-of-Plane Behaviour of As-Built …

5.6.6 Effect of Test Setup Throughout this experimental campaign, two different test setup were used to carry out the OOP tests, using airbags or pneumatic actuators. In both ones, distributed OOP loadings were applied to mobilize the entire panel. Due to that, it is important to compare the responses of the panels Inf_06 (no columns axial load, no previous IP damage, 10 mm plaster, OOP load applied with airbags) and Inf_08 (no columns axial load, no previous IP damage, 10 mm plaster, OOP load applied with airbags) and assess the variations among them due to the setup. Regarding the cracking pattern (Fig. 5.44a and b), trilinear cracking was observed in both specimens, however the spread cracking did not occur in Inf_08. From the force–displacement responses, shown in Fig. 5.44c, the differences, such related to the cracking point and the maximum peak load, become clear. Concerning the cracking strength, the specimen Inf_06 reached a value 2.73 times higher for a corresponding OOP drift 3.36 times higher. The maximum strength of the specimen Inf_06 was 77% higher than that of Inf_08 and occurred for an OOP drift 35% higher. Finally, the panel Inf_06 dissipated 28% higher energy than that if Inf_08. These results show again the high variability of the results concerning the OOP testing of panels built with hollow clay horizontal bricks.

5.7 Final Remarks and Major Observations from the Experimental Campaign An experimental characterization of the OOP behaviour of as-built infill walls was carried out, first resorting to quasi-static OOP tests with airbags and later with pneumatic actuators. Nine full-scale infill panels were tested with hollow clay horizontal bricks units 150 mm thick. The variables under analysis were the following ones: effect of columns axial load; effect of reduction of the panel support width; effect of the previous damage; effect of plaster, effect of test setup and effect of workmanship. From the comparative study, it was verified that: • The variables that affected most the OOP response of the specimens were the existence previous damage and the reduction of the panel support width and the existence of previous damage (due to IP demands); • It was observed that a panel subjected to a previous IP drift of 0.5% resulted in a reduction of the OOP maximum strength of 73% and consequently a reduction of the energy dissipation capacity around 91%; • The partial panel support width reduced the OOP maximum strength by about 59% and the energy dissipation capacity by about 70%. It should be also highlighted the importance of the test of the panel partially supported in the bottom beam, since it is a relevant contribution to the database presented in Chap. 3;

5.7 Final Remarks and Major Observations from the Experimental Campaign

221

Fig. 5.44 Effect of the test setup a Inf_06 cracking pattern; b Inf_08 cracking pattern; c Force– displacement curves; and d cumulative energy dissipation

• The workmanship seems to lead to important variability in terms of OOP maximum strength, failure mechanism and energy dissipation. The results revealed that this variation can be between 10 and 30%; • The plaster contributed to the slight increase of the OOP maximum strength around 3% and to the spread of the panel cracking. However, it was not observed any additional contribution in terms of energy dissipation; • Important variations were found due to the test setup in terms of maximum strength and energy dissipation capacity. Finally, it is important to highlight that the influence of the mortar strength properties may have had influence in the response of the panels, as well as other variables such as hardening time and experimental variability. Additional tests are needed to clarify the effect of each variable and, in particular, to assess the columns axial load effect in OOP response of the panels.

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References 1. Ani´c F, Penava D, Abrahamczyk L, Sarhosis V (2019) A review of experimental and analytical studies on the out-of-plane behaviour of masonry infilled frames. Bull Earthquake Eng 2. Furtado A, Costa C, Arêde A, Rodrigues H (2016) Geometric characterisation of Portuguese RC buildings with masonry infill walls. Eur J Environ Civil Eng:1–16 3. CEN (2000) NP EN206—concrete—specification, performance, production and conformity (Portuguese version). European Comittee for Standardization 4. CEN (2006) EN 196–2006 Methods of testing cement (European Committee for Standardization) 5. Griffith M, Vaculik J, Lam N, Wilson J, Lumantarna E (2007) Cyclic testing of unreinforced masonry walls in two-way bending. EarthqEng Struct Dyn 36:801–821 6. Ferreira T, Costa AA, Arêde A, Gomes A, Costa A (2015) Experimental characterization of the out-of-plane performance of regular stone masonry walls, including test setups and axial load influence. Bull Earthq Eng 13(9):2667–2692 7. NI (2012) National instruments. LabView software 8. Di Domenico M, Ricci P, Verderame GM (2019) Experimental assessment of the out-of-plane strength of URM infill walls with different slenderness and boundary conditions. Bull Earthq Eng 17(7):3959–3993 9. Ramos T, Furtado A, Eslami S, Alves S, Rodrigues H, Arêde A, Tavares PJ, Moreira PMGP (2015) 2D and 3D digital image correlation in civil engineering–measurements in a masonry wall. Proc Eng 114:215–222

Chapter 6

Experimental Assessment of Strengthening Solutions to Prevent the OOP Collapse of Infill Masonry Walls Through Textile Reinforced Mortars

6.1 Introduction Seismic events worldwide clearly showed that a crucial issue for life safety and loss reduction due to earthquakes for existing RC buildings is related to the OOP collapse of infill masonry walls. The adequate knowledge of all the aspects related to the behaviour of infilled framed structures, their components (structural and nonstructural elements) and their interaction is a fundamental issue to guide the practitioners in assessing and strengthening existing buildings. Therefore, the experimental validation of strengthening strategies available in the literature (presented and discussed in Chap. 3) and the development and testing of new possible strategies to reduce this seismic vulnerability and prevent the infill walls’ OOP collapse are nowadays paramount issues to be faced up. All the studies concerning the use of the TRM technique presented in Chap. 3, revealed interesting and promising results in terms of OOP strength increment and prevention of brittle failures, which should be enlarged and further analysed for future desirable design tools for practitioners. From, a technical point of view, all the techniques are effective if the strengthening material is well anchored to the surrounding frame and bonded to the panel. Different types of anchors were found in the literature, with different materials and different application procedures. Anyway, the number of experimental tests on the OOP retrofitted infill walls filling RC frames with TRM is still quite limited, especially if full-scale specimens are considered, and, due to the lack of data, specific design approaches are generally missing. Additionally, no seismic codes currently provide design tools for TRM strengthening for the considered application. This present chapter aims presenting an experimental campaign that was carried out to assess the efficiency of TRM-based solutions to improve infill panels OOP behaviour and prevent their collapse. Thus, the chapter is sub-divided into two major sub-sections, namely: (i) testing campaign of strengthened masonry wallets subjected

© The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 A. F. Furtado, Seismic Vulnerability Assessment and Retrofitting Strategies for Masonry Infilled Frame Building, Springer Theses, https://doi.org/10.1007/978-3-031-20372-5_6

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6 Experimental Assessment of Strengthening Solutions to Prevent …

to flexural strength tests, where the major details of the strengthening material application, test results and comparison with the as-built results are provided; and (ii) testing campaign of full-scale strengthened infill panels subjected to OOP loadings applied by pneumatic actuators with and without prior damage (due to previous IP test). The assessment of each strengthening solution is performed by comparing it with the as-built specimen and by assessing the impact of the prior damage on the performance of the strengthening technique.

6.2 Flexural Strength Tests in Strengthened Infill Masonry Wallets 6.2.1 Objectives and Specimens’ Description The strengthening of infill walls is an important topic nowadays, which needs further and deeper research towards developing design guidelines. These low-cost and fast solutions can be applied in rehabilitation interventions in existing buildings or new constructions. Therefore, the experimental testing here reported is of great importance since it can provide full insight regarding the response of the panels under seismic loadings. The aim of this sub-section is to assess the efficiency of TRM-based solutions on masonry wallets subjected to flexural strength tests parallel to the bed joints. A total amount of 20 masonry wallets, with the same geometric properties were strengthened with two different textile meshes, with and without connectors. The strengthening efficiency is assessed by comparing the results of the flexural strength capacity of the strengthened panels with the results reported in Sect. 4.3.5.1 regarding infill wallets with the same masonry units and geometric properties, without plaster and with plaster. Two different test setups were adopted to assess the flexural capacity of the strengthened panels, namely: 10 specimens (strengthened without connecters) were first subjected to the flexural strength test according to the standard NP EN 1052-2 [1], but during the testing campaign it was observed the same phenomena, already reported in Sect. 4.3.5.1, which related to the specimens shear failure due to the small geometric dimensions. Due to that, a second test setup was used to assess the panels’ flexural strength capacity by performing cantilever flexure strength tests. In this setup, 10 specimens (strengthened with connecters) were tested up to the failure. Throughout this sub-section, it is first presented the description of the infill wallets, the adopted strengthening strategies as well the details concerning their strengthening process. After that, details concerning the test setup and instrumentation used in the cantilever flexure strength tests are provided. The details concerning the test setup and instrumentation used in the flexural strength tests according to the standard NP EN 1052-2 [1] were previously reported in Sect. 4.3.5.1. The experimental results are presented and discussed in terms of force–displacement curves and specimens’

6.2 Flexural Strength Tests in Strengthened Infill Masonry Wallets

225

failure mode. The efficiency of the strengthening techniques is assessed by comparing the results with those obtained in the as-built reference ones, reported in Sect. 4.3.5.1.

6.2.2 Specimens Description and Strengthening Process The twenty infill wallets were built with the same geometric properties of those tested as described in Sect. 4.3.5.1, since it was aimed at comparing them to assess the efficiency of the strengthening solutions. Thus, the infill wallets were built with 600 mm width, 620 mm height (due to the use of 10 mm plaster on the top and bottom of the specimens and 150 mm thickness (in its as-built condition). The strengthening was applied to only one of their faces. The twenty infill wallets were divided into two groups according to the textile mesh used, namely 10 specimens strengthened using a GFRP textile mesh and 10 specimens strengthened using a polypropylene (PP) textile mesh. In each group, 5 were strengthened without panel-textile mesh connectors and 5 with connectors. GFRP strengthening The GFRP textile mesh is herein considered a “strong” mesh, supplied by FASSA BORTOLO, designated FASSA NET ARG 40, with tensile strength equal to 70kN/m, mesh size equal to 40 × 40 mm, maximum extension at rupture equal to 3%, mesh density equal to 315 g/m2 and is a bidirectional one, as shown in Fig. 6.1a. Metallic connectors (φ 6 mm and 8 cm length, Fig. 6.1b) ensured the connection of the mesh to the panel and that the connectors crossed two of the brick internal septa. Four metallic connectors were applied to each of the five of the specimens. The schematic layouts of the strengthening strategy adopted for the specimens without and with connectors are shown in Fig. 6.1c and d. The strengthening process followed the following steps, namely: . After the infill wallets construction, they were mortar splashed and after 28 days they were wet with water, as shown in Fig. 2a; . Application of the first layer of plaster 5 mm thick, as shown in Fig. 6.2b; . Application of textile mesh and second layer of plaster 5 mm thick, as shown in Fig. 6.2c; . As can be observed in Fig. 6.2c, it was placed a textile mesh with a total length equal to 1050 mm, which means a total of 600 mm for the panel height and remaining 450 mm for each top and bottom of the panel (150 mm for the top of the panel and 300 mm for the panel back). The GFRP mesh was folded to envelope the panel, and it was anchored in the top and back side of the panel with 3 metallic connectors each, as shown in Fig. 6.2d and e. After that, the 5 mm plaster thickness was applied to cover the textile mesh on the top and back side of the panel, as shown in Fig. 6.2f.

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6 Experimental Assessment of Strengthening Solutions to Prevent …

a)

b)

600mm

600mm

620mm

620mm

400mm

400mm

Panel Plaster

GFRP mesh c)

Metallic connector Panel GFRP mesh

Plaster d)

Fig. 6.1 Strengthening materials used for the GFRP group specimens: a detail of the textile mesh; b metallic connectors; c strenghtening strategy adopted for specimens without connectors; and d strenghtening strategy adopted for specimens with connectors

PP strengthening The PP textile mesh is herein considered a “weak” mesh, a “Tenax aviary” type mesh, with tensile strength equal to 5.25 kN/m, mesh size equal to 16 × 19 mm, maximum yielding extension equal to 15%, mesh density equal to 315 g/m2 and is bidirectional one, as shown in Fig. 6.3a. The same metallic connectors were used in the specimens of the group PP. The schematic layout of the strengthening strategy adopted for the specimens without and with connectors are shown in Fig. 6.3b and c. The same strengthening procedure used in the GFRP group was adopted in the PP specimens. The mortar used to build the infill panels and used for the plaster was again a M5 class (“Ciarga”), as referred in Chaps. 4 and 5. The tensile strength of the mortar is 1.56 MPa with a C.o.V. equal to 5.4% and a standard deviation 0.22 MPa. The compressive strength is 4.09 MPa with a C.o.V. equal to 6.9% and a standard deviation equal to 0.11 MPa.

6.2 Flexural Strength Tests in Strengthened Infill Masonry Wallets

227

a)

b)

c)

d)

e)

f)

Fig. 6.2 Strengthening process of the GFRP group specimens: a splashing of the panels; b application of the first layer of the plaster (about 5 mm); c application of the textile mesh and second layer of plaster (5 mm); d fixation of the mesh to the top of the panel; e lateral view of the metallic connectors fixing the mesh to the top and backside; and f application of 5 mm thick plaster

6.2.3 Test Setup Description and Instrumentation The test setup 2 was adopted aiming at avoiding the problems observed in the specimens tested with the test setup 1, which will be described in the next sub-sections. Cantilever flexure strength tests were carried out in the remaining 10 strengthened

228

6 Experimental Assessment of Strengthening Solutions to Prevent …

a)

600mm

620mm

PP mesh

Plaster b)

Panel c)

Fig. 6.3 Strengthening materials used for the PP group specimens: a detail of the textile mesh; b strenghtening strategy adopted for specimens without connectors; and c strenghtening strategy adopted for specimens with connectors

specimens with connectors. Before performing the test, it was mandatory to carry out a preliminary and auxiliary preparation of all the specimens, namely by filling a half-brick with mortar, as shown in Fig. 6.4. The aim of this procedure was to provide the specimen with a bottom part more stiff and strong to be anchored to a steel profile. With this, it was possible to bend the specimen without crushing of the anchored base and thus to push the mesh up to the tensile failure as desired. A vertical servo-hydraulic actuator applied the cantilever load with a maximum capacity equal to 100kN (±100 mm). The load perpendicular to the infill panels was applied in the top of the infill panel laid in a horizontal plan, as shown in Fig. 6.6. In the extremity of the actuator, a metallic hinge was included that allowed to continuously apply the OOP load during the panel bending and rolling during the test. A steel plate was included between the specimen and the panel to distribute the transverse load over the top of the panel. The bottom part of the panel (half-height brick that was reinforced with mortar) was anchored to a steel HEB300 shape, by a quadrangular steel shape to distribute the anchorage along 10 cm width of the panel and was attached to the base steel shapes by two steel rods ø20mm. The schematic layout of the test setup is shown in Fig. 6.5. The test setup general view is shown in

6.2 Flexural Strength Tests in Strengthened Infill Masonry Wallets

a)

b)

c)

d)

229

Fig. 6.4 Cantilever flexure strength tests: preparation of the specimens: a overview of the specimens; b filling with mortar process; c top view of half-brick filled with mortar; and d profile view of the final aspect

Fig. 6.6, where some details concerning the application of the load (Fig. 6.6b), and panel anchorage (Fig. 6.6c) can be found. Regarding the specimens’ instrumentation, it was used a total of four LVDT to measure the panel OOP displacements and one clinometer (to record the panel rotation along the flexure direction) that was attached to the bottom anchored part of the panel to verify it occur some rotation during the test. The schematic layout of the instrumentation adopted is shown in Fig. 6.7a. The detail of the clinometer position is shown in Fig. 6.7b.

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6 Experimental Assessment of Strengthening Solutions to Prevent …

Fig. 6.5 Cantilever flexural strength tests: test setup: schematic layout

a)

b)

c)

d)

Fig. 6.6 Cantilever flexural strength tests: test setup: a left view; b right view; c detail of the panel anchorage; and d general overview

6.2 Flexural Strength Tests in Strengthened Infill Masonry Wallets

231

50mm

250mm

LVDTs LVDTs

100mm 100mm 100mm

Clinometer a)

b)

Fig. 6.7 Cantilever flexural strength tests: instrumentation: a schematic layout b clinometer view

6.2.4 Experimental Results The experimental results are presented according to each test setup by analysing the OOP force–displacement curves and the most representative failure modes. Afterwards, it is assessed the maximum flexure strength of each strengthening solution to perform the comparison between the results obtained in both test setups and to analyse the effect of the connectors. Finally, the efficiency of the strengthening strategies is analysed by comparing the results of the strengthened panels with the as-built results. The flexural strength of the specimens subjected to cantilever flexure tests was computed according to Eq. 6.1, where F is the OOP load, hpanel and wpanel are the panel height and width, respectively, and t is the panel thickness. f b,paralell,2 =

F × h panel × 6 wpanel × t 2

(6.1)

The force–displacement curves of the GFRP and PP group specimens (strengthened with connectors and tested with the test setup 2 are shown in Fig. 6.8a and b. From the results the following observations can be drawn, namely: . The maximum force reached by the GFRP group was 3.62kN and the lowest was 2.58 kN. The average value was equal to 3.15 kN with a C.o.V. equal to 12%, which is significantly lower than the variation of the results of four point flexure strength tests; . The range of the OOP displacements at the maximum OOP load was between 28.30 mm and 60.85 mm. A mean value equal to 45.33 mm and a C.o.V. equal to 29% was found. Concerning the mean ultimate displacement, it was about 78 mm with a C.o.V. equal to 19%, as summarized in Table 6.1

6 Experimental Assessment of Strengthening Solutions to Prevent … GFRP strengthening with connectors

FOOP(kN)

4.5

2.4

1.8

1.2

1.5

0

20

40

60

80

Specimen 1 Specimen 2 Specimen 3 Specimen 4 Specimen 5 µ curve µ−2SD curve µ+2SD curve

Setup 2 Textile mesh: PP With connectors

3.0

0.0

6.0

4.5

FOOP(kN)

Specimen 1 Specimen 2 Specimen 3 Specimen 4 Specimen 5 µ curve µ−2SD curve µ+2SD curve

Setup 2 Textile mesh: GFRP With connectors

MOOP (kN.m)

6.0

PP strengthening with connectors

1.2

0.6

1.5

0.6

0.0 100

0.0 0

20

40

0.0 100

80

b) 0.8

Specimen 1 Specimen 2 Specimen 3 Specimen 4 Specimen 5 µ curve µ−2SD curve µ+2SD curve

Setup 2 Textile mesh: GFRP With connectors

0.6

Specimen 1 Specimen 2 Specimen 3 Specimen 4 Specimen 5 µ curve µ−2SD curve µ+2SD curve

Setup 2 Textile mesh: PP With connectors

0.7 0.6

0.5

fb,parallel (MPa)

fb,parallel (MPa)

60

OOP displacementaverage (mm)

a)

0.7

1.8

3.0

OOP displacementaverage (mm)

0.8

2.4

MOOP(kN.m)

232

0.4 0.3

0.5 0.4 0.3

0.2

0.2

0.1

0.1 0.0

0.0 0

20

40

60

80

100

120

0

20

OOP displacementaverage (mm)

40

60

80

100

120

OOP displacementaverage (mm)

c)

d)

Fig. 6.8 Cantilever flexural strength tests: force–displacement curves a GFRP group b PP specimens; flexural strength c GFRP group; and d PP group

Table 6.1 Summary results of the cantilever flexural strength tests: GFRP group Group GFRP

Specimens

FOOP,max (kN)

dF,OOP,max (mm)

fb,parallel,max (MPa)

dF,OOP,ult (mm)

Specimen 1

2.58

28.30

0.45

56.01

Specimen 2

3.25

60.85

0.57

85.35

Specimen 3

3.62

55.45

0.64

75.70

Specimen 4

3.26

44.85

0.57

76.80

Specimen 5

3.03

37.20

0.53

97.10

Average

3.15

45.33

0.55

78.19

12.12

29.18

12.58

19.28

0.38

13.23

0.07

15.07

C.o.V. (%) SD

. The flexural strength of the GFRP group specimens varied between 0.45 MPa and 0.64 MPa, with an average value equal to 0.55 MPa and a C.o.V. equal to 12%, as shown in Fig. 6.8c;

6.2 Flexural Strength Tests in Strengthened Infill Masonry Wallets

233

. Concerning to the specimens of PP group, it can be observed that the maximum OOP force varied between 1.20 kN and 1.89 kN, with an average value equal to 1.52 kN and a corresponding C.o.V. around 16%; . The OOP displacement corresponding to the maximum OOP force varied slightly between 8.85 mm and 12.65 mm, with an average value equal to 11.73 mm. Concerning to the ultimate OOP displacement, an average value equal to 31.14 mm was obtained with a C.o.V. around 12%. The flexural strength of the specimens from PP group varied between 0.21 MPa and 0.33 MPa, as summarized in Table 6.6, and shown in Fig. 6.8d. An average value equal to 0.27 MPa was found. Regarding the comparison between the results obtained by each group, it can be observed that the flexural strength reached by the GFRP group was two times higher with a C.o.V. around 17%. A big difference was found relatively to the OOP displacement at the maximum load, where the GFRP group reached an average value of 45.33 mm, 3.86 times higher than the PP group. Regarding the ultimate OOP displacement, it can be observed that the GFRP reached an average value equal to 78.19 mm, 2.51 times higher than the PP group. These results are quite different from those reached with test setup 1. The shear failure was avoided with this test setup, which allows for better understanding the specimens’ flexural behaviour. Nonetheless, complementary tests should be performed in specimens with larger geometry to validate the results herein presented (Table 6.2). As expected, the OOP displacements reached in the cantilever flexure strength test are larger than the ones reached with the four point flexure strength tests, which is related to the fact that in the test setup 1, the cantilever has total span length equal to L. On the other hand, the span length of the test setup 2 is equal to L/3. Concerning the failure modes of the specimens herein presented, as previously mentioned, the shear failure was prevented and all the ruptures were due to flexure. Starting from the failure mode of the GFRP group, shown in Fig. 6.9, they were characterized by the development of horizontal cracks (Fig. 6.9a), and the high deformation capacity of the panel without collapse (Fig. 6.9b) with some crushing of the back side of the wallets due to stresses concentrations. Regarding the PP specimens Table 6.2 Summary results of the cantilever flexural strength tests: PP group Group

Specimens

PP

Specimen 1

1.63

11.85

0.29

27.30

Specimen 2

1.20

12.65

0.21

36.30

Specimen 3

1.40

12.60

0.25

31.10

Specimen 4

1.50

8.85

0.26

27.70

Specimen 5

1.89

12.70

0.33

33.30

Average C.o.V. (%) SD

FOOP,max (kN)

dF,OOP,max (mm)

fb,parallel,max (MPa)

dF,OOP,ult (mm)

1.52

11.73

0.27

31.14

16.92

14.04

16.77

12.21

0.26

1.65

0.04

3.80

234

6 Experimental Assessment of Strengthening Solutions to Prevent …

a)

b)

Fig. 6.9 Cantilever flexural strength tests: GFRP group failure modes a development of flexural crack b top view of the damaged specimen

failure, they were all also due to flexure with development of horizontal cracks, as shown in Fig. 6.10a, followed by the collapse of the panel for larger OOP displacements (Fig. 6.10b). It was possible to observe the failure of the mesh due to high tensile demands (Fig. 6.10c and d).

6.3 Experimental Assessment of TRM Solutions to Improve the Out-of-Plane Capacity of Full-Scale Infill Masonry Walls 6.3.1 Objectives and Specimens’ Description All the studies mentioned in Sect. 3.2.4 about the TRM technique revealed interesting and promising results in terms of increasing the OOP strength of the panels and preventing brittle failures and collapse, which should be enlarged and further analysed for possible development of desirable design tools for practitioners. From, a technical point of view, all the solutions are likely to be effective if the strengthening material is well anchored to the surrounding frame and bonded to the panel. Different types of anchors were found in the literature, with different materials and different application procedures. Anyway, the number of OOP experimental tests of infill walls strengthened with TRM is still quite limited, especially full-scale specimens. Additionally, no seismic codes currently provide design tools for TRM strengthening. This section presents the experimental analysis of TRM-based strengthening solutions to prevent the out-of-plane collapse of masonry infills in RC buildings. The experimental campaign comprises the OOP testing of five full-scale infill walls made up of horizontal hollow clay bricks, 150 mm thick, two of them unstrengthened and

6.3 Experimental Assessment of TRM Solutions to Improve …

a)

d)

235

b)

e)

Fig. 6.10 Cantilever flexural strength tests: PP group failure modes a development of horizontal flexural crack b collapse of the panel; c lateral view of the panel collapsed; and d detail view of the mesh failure

the remaining three strengthened, with and without prior damage. All the specimens have the same geometric characteristics of those tested and presented in Chap. 5. In this testing campaign, the reference specimens are the panels Inf_08 and Inf_11, and the strengthened panels are Inf_10, Inf_12 and Inf_13. The following considerations (by each specimen) can be added: . Inf_08: One-leaf panel; aligned with the external face of the supporting beam; OOP half-cyclic load, no columns axial load; no previous damage; total panel width supported in the bottom beam; 10 mm plaster; OOP load applied with pneumatic actuators; . Inf_11: One-leaf panel; aligned with the external face of the supporting beam; OOP half-cyclic load, no columns axial load; with previous damage due to IP test (0.3% inter-storey drift); total panel width supported in the bottom beam; 10 mm plaster; OOP load applied with pneumatic actuators; . The remaining three specimens (Inf_10, Inf_12 and Inf_13) were strengthened to prevent the out-of-plane collapse with two strengthening strategies, all based on the application of systems made up of common mortar plaster and glass fibrereinforced polymer meshes. The influence of the mesh properties (tensile strength

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6 Experimental Assessment of Strengthening Solutions to Prevent …

and ductility) and the efficiency of the strengthening under combined IP-OOP loadings are investigated in this experimental testing campaign. Details regarding the strengthened specimens will be provided in Sect. 6.3.2. All the tests consisted in the application of a semi-cyclic (loading–unloadingreloading) history of imposed displacements in the OOP direction by means of a uniformly distributed load resorting to the system detailed in Sect. 5.5.1. The mechanical properties of the adopted materials are characterized and presented. The results are presented in terms of OOP force–displacement responses, deformed shapes, damage evolution, energy dissipation capacity and damping. In the end, the tests’ results are compared to each other to assess the effectiveness of the selected strengthening techniques and to provide a support towards the choice of the best strategies for future further investigations and applications.

6.3.2 Description of the Strengthening Strategies All the specimens were strengthened by using a fibreglass (GFRP) reinforcing mesh coupled to non-ductile (hereinafter referred to as “common”) cement-mortar. Table 6.3 reports a summary of the five specimens tested and analysed in the following sections. The characteristics of each specimens are the following: . Inf_10: strong mesh, plastic connector to anchor the mesh-infill panel; steel connector to anchor the mesh to the frame, common cement-mortar M5 class; . Inf_12: medium mesh; plastic connector to anchor the mesh-infill panel; steel connector to anchor the mesh to the frame plus a thin steel plate distributed along the contour where the connectors were applied; common cement-mortar M5 class; . Inf_13: Same solution of the specimen Inf_12. 6.3.2.1

Specimen Inf_10

The specimen Inf_10 was strengthened using a TRM solution (herein designated strengthening solution A) with the glass fibre mesh “FASSANET ARG 40”, supplied by Fassa Bortolo, with a matrix 4 × 4 cm and a tensile strength equal to 70 kN/m. Table 6.3 Experimental tests matrix Specimens

Strengthening

Type of mesh

Previous IP damage

Note

Inf_08

No

N/A

No

Reference specimen

Inf_10

Yes

Strong

No

Comparison with Inf_08

Inf_11

No

N/A

Yes (0.3%)

Reference specimen

Inf_12

Yes

Medium

Yes (0.3%)

Comparison with Inf_11

Inf_13

Yes

Medium

No

Comparison with Inf_08

6.3 Experimental Assessment of TRM Solutions to Improve …

237

This mesh was the same used in the wallets subjected to flexural strength tests and described in Sect. 6.2.2. Steel connectors (φ6mm and 8 cm length) ensured the connection of the mesh to the panel and that two of the brick septa were crossed by (Fig. 6.11a and b). In order to ensure a better connection of the mesh to the wall, plastic discs with 6 cm diameter were used with the metallic connectors. Regarding the anchorage of the mesh to the RC frame it was adopted a solution made with M8 steel connectors also with a plastic disk with 6 cm diameter, as shown in Fig. 6.11c and d. The strengthening application followed the steps described hereafter: . Application of a first layer of plaster (thickness around 0.5 cm), shown in Fig. 6.12a; . Placement and positioning of the mesh, shown in Fig. 6.12b; . Fixation of the mesh with the connectors, shown in Fig. 6.12c; . Application of a second layer of plaster 1–1.5 cm thick, shown in Fig. 6.12d.

a)

b)

c)

d)

Fig. 6.11 Strengthening materials used for the panel Inf_10: a detail of the steel connectors and plastic disk used in the panel; b general view of the mesh fixed to the panel; c detail of the M8 steel connectors and plastic disk used to fix the mesh to the frame; and d detail view of the application

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6 Experimental Assessment of Strengthening Solutions to Prevent …

a)

b)

c)

d)

Fig. 6.12 Strengthening process of the panel Inf_10: a Application of the first layer of plaster; b positioning of the mesh; c anchorage of the mesh to the frame; and d application of the second layer of mortar

As previously mentioned, the mortar used for the plaster application was a current one, M5 class. Five vertical strips with 1 m width each were applied with an overlapping width equal to 10 cm, according to the supplier recommendation. The transition between the panel and the frame elements was reinforced with two layers of mesh with a total width equal to 30 cm (15 cm in the panel plus 15 cm in the frame elements). The schematic layout of the strengthening strategy adopted is plotted in Fig. 6.13.

6.3.2.2

Specimen Inf_12 and Inf_13

It was adopted the same strengthening solution for the panels Inf_12 and Inf_13, again a TRM solution and herein designated solution B. The used reinforcing mesh was a bidirectional fibreglass (GFRP) mesh with a 16.7 × 16.7 mm2 matrix (see Fig. 6.14a), a weight of 185 g/m2 , a nominal tensile strength equal to 40.0 kN/m, and a maximum ultimate strain equal to 3.4%.

6.3 Experimental Assessment of TRM Solutions to Improve …

239

INFILL PANEL

Fig. 6.13 Schematic layout of the Inf_10 strengthening strategy

The application procedure of this strengthening strategy started with the application of 1 cm plaster. Then plastic connectors, see Fig. 6.14b, were applied throughout the infill panel to position and fix the reinforcing mesh. The plastic connectors have been applied for the tested specimens to position and fix the mesh. The roll of mesh was provided with 1 m width and 50 m length. Five vertical strips (1 m width) were used to strengthen the wall overlapped to each other. The application of vertical strips resulted easier with respect to the application of horizontal strips (whose length can also vary depending on the bay length). The overlap length used between each vertical strip was 10 cm. The mesh was extended for 15 cm both on the beams and columns, Then, in the overlapping regions for the transition RC frame-infill panel it was assumed a duplicated mesh with an overlap equal to 30 cm (15 cm for the RC frame and 15 cm for the infill panel). A different strategy was adopted for the mesh-frame anchorage, namely it was adopted a combined solution using a steel plate (3 mm thick and 30 mm width) along the alignment defined for the mesh-frame anchors. Ø 10 mm holes were drilled in each location defined for the anchor to insert M8 steel connectors. Thus, the mesh was placed between the frame and the steel plate, which was anchored with the steel connectors to the frame. The main goal was to minimize the local sliding/shear failure of the mesh, a common problem reported by [2]. Details of the steel connectors and the steel plate are shown in Fig. 6.14c and d.

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6 Experimental Assessment of Strengthening Solutions to Prevent …

Fig. 6.14 Strengthening materials used for the panel Inf_12 and Inf_13: a general view of the mesh fixed to the panel; b detail of the plastic connector for the mesh-panel anchorage; c detail of the M8 steel connectors for the frame-mesh anchorage and general view of the steel plate; and d detail view of the steel conector and plate

The strengthening procedures for the panels Inf_12 and Inf_13 were according to the following steps: . Application of a first layer of plaster (thickness around 0.5–1 cm), shown in Fig. 6.15b; . Placement and positioning of the mesh, shown in Fig. 6.15c; . Fixation of the mesh to the top RC beam, with the application of the steel plate and connectors, shown in Fig. 6.15c. Details of the holes’ drilling in the frame for the connector are shown in Fig. 6.15d. The tightening of the steel connector is shown in Fig. 6.15e. The detail of the mesh-frame anchorage is shown in Fig. 6.15f; . Application of the plastic connectors for the mesh-panel anchorage. The general view of the specimen with the plastic connectors is shown in Fig. 6.15g; . Application of a second layer of plaster 2 cm thick, shown in Fig. 6.12h. The final layout of the strengthening solution herein designated “B” is shown in Fig. 6.16.

6.3 Experimental Assessment of TRM Solutions to Improve …

241

a)

b)

c)

d)

e)

f)

g)

h)

Fig. 6.15 Strengthening process of the panels Inf_12 and Inf_13: a general view of the panel asbuilt condition; b application of the first layer of plaster; c positioning of the mesh; d holes drilling in the column for connector; e tightening of the steel connector; f detail view of the mesh-frame anchorage; g general view of the strengthened panel with mesh and all the connectors applied; and h application of the second layer of mortar

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6 Experimental Assessment of Strengthening Solutions to Prevent …

(a)

(b) Fig. 6.16 Strengthening process of the panel Inf_12: a Schematic layout; and b general view of the strengthened specimen

6.3.3 Material Properties The properties of the mortar used for strengthening (compressive and tensile strength) were evaluated according EN 196-2006 [3] standard. Each value of compressive strength and tensile strength shown in Table 6.4 is evaluated based on six prismatic samples. From the results, it is possible to observe that the compressive strength of the mortar used for the infill bed joints are around 5 MPa as expected, being the highest

6.3 Experimental Assessment of TRM Solutions to Improve …

243

Table 6.4 Results from flexure and compressive strength tests on mortar specimens Mortar for infill bed joints

Mortar for strengthening/plaster

Flexure strength (MPa)

Compressive strength (MPa)

Flexure strength (MPa)

Compressive strength (MPa)

Inf_08

1.90 (C.o.V. = 6.6%)

5.24 (C.o.V. = 7.8%)

1.81a (C.o.V. = 7.3%)

4.01a (C.o.V. = 8.2%)

Inf_10

1.82 (C.o.V. = 11.6%)

4.74 (C.o.V. = 10.9%)

2.28 (C.o.V. = 5.8%)

4.94 (C.o.V. = 7.2%)

Inf_11

2.22 (C.o.V. = 10.7%)

6.49 (C.o.V. = 11.3%)

2.32a (C.o.V. = 4.9%)

8.20a (C.o.V. = 5.5%)

Inf_12

1.84 (C.o.V. = 7.6%)

6.25 (C.o.V. = 8.9%)

1.65 (C.o.V. = 7.8%)

5.71 (C.o.V. = 5.8%)

Inf_13

1.53 (C.o.V. = 4.9%)

3.52 (C.o.V. = 6.6%)

2.56 (C.o.V. = 11.9%)

6.52 (C.o.V. = 8.6%)

Specimen ID

a

Mortar properties of the plaster

value achieved by the specimen Inf_11 with 6.49 MPa and the lowest by Inf_13 with 3.52 MPa. Concerning the flexural strength, the values are around 1.8 MPa, being the largest value obtained again by Inf_11 with 2.22 MPa and the lowest by Inf_13 with 1.53 MPa. The results of the mortar used for the strengthening/plaster present similar trends, being the highest values achieved for the specimen Inf_11 and the lowest for Inf_08.

6.3.4 Test Setup, Instrumentation and Loading Protocol All the specimens presented herein were subjected to quasi-static tests with pneumatic actuators using the same test setup described in Sect. 5.5.1, as shown in Fig. 102. It was assumed the instrumentation layout described in Sect. 5.5.2 in all specimens test. Therefore, the applied loading protocol again consisted of halfcyclic OOP displacements (loading–unloading-reloading) that were imposed with increasing displacement levels, twice for each dis-placement level, following the protocol described in Sect. 5.5.2.

6.3.5 Experimental Results In this section, the results of the experimental testing procedure described above are presented and analysed test by test responses in terms of OOP load (FOOP ) versus OOP displacement in the centre of the infill panel (dOOP ) are shown, together with the damage evolution from the first observed cracks on the panel until the end of the test. The results of the reference specimens Inf_08 and Inf_11 will not be hereafter

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6 Experimental Assessment of Strengthening Solutions to Prevent …

presented, since they were detailed and discussed in Sects. 5.5.3.1 and 5.5.3.3, respectively. Only the results of the strengthened specimens Inf_10, Inf_12 and Inf_13 are described in the following sub-sections.

6.3.5.1

Specimen Inf_10 (Strengthening Solution A; Without Prior IP Damage)

The strengthened panel Inf_10 was only subjected to a pure OOP test, similarly to the reference specimen Inf_08. The force–displacement curve is shown in Fig. 6.17a. The initial (secant) stiffness of the FOOP − dOOP response, calculated as the ratio between FOOP and dOOP at the first peak relative to the first applied displacement level, is equal to kOOP,sec,ini = 18.90 kN/mm. Afterwards, the first macro-cracking occurred for an OOP drift equal to 0.33% and a corresponding force equal to 47.23 kN. The first cracking was a horizontal crack located at the panel mid-height with a total length corresponding to 2/3 of the panel length, as shown in Fig. 6.17b. Diagonal cracks and a parallel horizontal crack developed with the increase of the OOP displacements. The maximum peak load equal to 77.10 kN was reached for an OOP drift equal to 2.35%. After this point, it was visible the beginning of strength degradation. The displacement transducers were not capable to capture the maximum value of that displacement (which, however and based on the remaining transducers, was possible to estimate around 50 mm), because they exceeded their measuring range. However, it was possible to measure the value of the residual displacement of the wall (about 41.9 mm) and to restart the test from that point (Fig. 6.17a). The force–displacement response includes a hidden red line at the moment from which it was not possible to record the OOP displacements. After that moment, the stiffness degradation was more evident, but still ten more cycles were possible to be performed until reaching the ultimate drift equal to 6.31% without any strength degradation. The ultimate strength was 39.40 kN. The final damage of the panel is shown in Fig. 6.17c, and was characterized by a distributed cracking, particularly from the panel mid-height to the bottom. Figure 6.18a shows the lateral view of the bottom part of the panel where no panel detachment or sliding is visible. Some concentration of horizontal cracks are noticed, which are related to the failure of the bricks from the first row of the panel, as evidenced in Fig. 6.18b and c that showing the nearside of the panel. Some crushing of bricks was also visible as well as some sliding of the panel from bottom beam. It was observed that the failure of those bricks occurred at the maximum peak load and caused the large decrease of OOP strength and the larger OOP displacements.

6.3.5.2

Specimen Inf_12 (Strengthening Solution B, with Prior IP Damage)

The specimen Inf_12 was subjected to an IP-OOP test sequence, i.e. the panel was first subjected to a quasi-static IP test with a maximum drift equal to 0.3% and

6.3 Experimental Assessment of TRM Solutions to Improve …

245

Out-of-plane drift at control node (%) 0.00 100

3.48

5.22

6.96

8.70

10.43

Specimen Inf_10 No previous In-Plane drift Strengthening solution: A

80

Out-of-plane force FOOP (kN)

1.74

Displacements not recorder (out of range)

60 Ultimate stage

40

20

Cracking point Peak Load 0 0

20

40

60

80

100

120

Out-of-plane displacement at control node dOOP (mm)

a)

b)

c)

Fig. 6.17 Experimental results: Specimen Inf_10: a force–displacement curve; b first cracking; and c cracking pattern

afterwards to the OOP test, similarly to the reference specimen Inf_11. During the IP, slight detachment occurred between the panel and top beam, as proved by the horizontal crack in the top, evidenced in Fig. 6.19. Some plaster detachment occurred, however without any shear sliding failure of the mesh. A diagonal crack from the right top corner of the panel to the left bottom corner developed for the maximum IP drift (0.3%). Horizontal and vertical cracks over the panel contour were noticed with some concentration in the panel corners, due to the increase of the panel stresses during the distortion of the frame. The force–displacement curve is plotted in Fig. 6.20a, where it is possible to observe that the initial secant OOP stiffness of the panel was equal to 7.44 kN/mm and the first cracking occurred for an OOP drift equal to 0.71% and a corresponding force equal to 44.90 kN. The observed crack was diagonal, parallel to that developed during the IP test, as shown in Fig. 6.20b. After the first cracking, some diagonal cracks and horizontal ones occurred from all over the panel until the maximum peak load equal to 75 kN corresponding to an OOP drift equal to 2.75%.

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6 Experimental Assessment of Strengthening Solutions to Prevent …

a)

b)

c)

Fig. 6.18 Experimental results: Specimen Inf_10: a Lateral view of the damage in the front of the panel; b right view of the panel near side; and c left view of the panel back side

a)

b)

Fig. 6.19 Experimental results: Specimen Inf_12: a general view of the final damage after IP test; and b cracking pattern after IP test

At the peak load, the panel detached from the bottom beam, but with some sliding of the mesh from the steel plate anchored to the beam. The panel OOP strength reduced 20% until 60kN, and continued to reduce the strength until the maximum drift equal to 8.91% with a residual force equal to 39.16 kN. The final damages and cracking pattern of the specimen Inf_10 are shown in Fig. 6.20c. From the test, the strengthening solution was found effective since it prevented OOP collapse of the panel. During the test, large OOP displacements occurred in the bottom of the panel, which again were due to the failure of bricks distributed in the panel first row. Figure 6.21a presents the lateral view of the bottom part of the wall where the steel plate is still anchored to the beam, but exhibiting sliding failure of the anchored mesh at mid-length. An OOP displacement of the panel equal to 120 mm

6.3 Experimental Assessment of TRM Solutions to Improve …

247

Fig. 6.20 Experimental results: Specimen Inf_12: a force–displacement curve; b first cracking; and c cracking pattern

was measured, i.e. 70% of the panel thickness, which revealed the impressive capacity of the mesh to sustain the panel, as shown in Fig. 6.21b. The near side of the panel shows crushing of some bricks located in the top and bottom rows which can be justified by the fact that they were subjected to larger flexural than their flexural strengths (Fig. 6.21c and d). Figure 6.21e and f shows the nearside of the panel, where it is possible to observe the panel deformation. In fact, the panel was almost totally detached from the bottom support beam without collapsing. There was no detachment or sliding failure of the mesh in the other regions of the panel.

6.3.5.3

Specimen Inf_13 (Strengthening Solution B, Without Prior IP Damage)

The strengthened panel Inf_13 was only subjected to a pure OOP test, similarly to the reference specimen Inf_08, and to assess the efficiency of the strengthening solution B first implemented and tested in specimen Inf_12. The force–displacement curve

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6 Experimental Assessment of Strengthening Solutions to Prevent …

a)

b)

c)

d)

e)

f)

Fig. 6.21 Experimental results: Specimen Inf_12: a detail of the mesh sliding; b detail of the detachment of the panel and the mesh sliding; c left view of the panel back side; d right view of the panel back view; e detail of the deformed shape of the panel; and f detail of the measurement of the OOP displacement

is shown in Fig. 6.22a. The initial secant stiffness was equal to 80.92 kN/mm, as visible in the quick increase of the OOP strength for low displacements. The first macro-cracking occurred for an OOP drift equal to 0.17% and was an horizontal crack located above the panel mid-height, as shown in Fig. 6.22a. Again, diagonal and horizontal cracks appear until the maximum peak load (92.31 kN) corresponding to a drift of 2.55%. After that, it occurred the fracture/crushing of the

6.3 Experimental Assessment of TRM Solutions to Improve …

249

Fig. 6.22 Experimental results: Specimen Inf_13: a force–displacement curve; b first cracking; and c cracking pattern

bricks located at the first and, consequently, the OOP strength reduced about 44 kN and the OOP drift suddenly increased to approximately 7%. Until the last stage of the test, it was not observed any significant reduction of the panel strength and the residual capacity of the panel was found equal to 42.61 kN for an OOP drift equal to 8.70%. The test has stopped due to the limit of the reached pneumatic actuator stroke. The panel cracking pattern is shown in Fig. 6.22c. After the test, a detailed damage assessment was carried out. The main focus was to assess the anchorage of the mesh, the performance of the frame-mesh connectors, and the importance of the steel plate. This test evidenced the importance of the adopted frame-mesh anchorage since it allowed to reach high OOP displacements avoiding the shear sliding of the mesh. Figure 6.23a shows the lateral view of the bottom part of the panel where it possible to observe the mesh-frame anchorage. The superficial plaster layer was removed to verify the state of the mesh, shown in Fig. 6.23b. The mesh was found in good conditions, with one example of mesh sliding illustrated in Fig. 6.23c.

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6 Experimental Assessment of Strengthening Solutions to Prevent …

a)

b)

c)

d)

e)

f)

Fig. 6.23 Experimental results: Specimen Inf_13: a lateral view of the bottom part of the panel front; b detail of the mesh-frame anchorage after detached the plaster; c detail of few examples of sliding of the mesh; d detail of the top mesh-frame anchorage; e left view of the panel back side; and f right view of the panel back side

The same procedure was performed in the top part of the wall, where it was possible to observe that the mesh was intact with perfect anchorage to the frame (Fig. 6.23d). As previously described, the back side of the panel (Fig. 6.23e and f) shows the fracture and crushing of the bricks located in the first and last row of the panel, in particular in the mid-length zone.

6.3 Experimental Assessment of TRM Solutions to Improve …

251

6.3.6 Comparison of the Results and Discussion In this section, the experimental results above described are compared to each other in terms of OOP force–displacement envelopes (Sect. 6.3.6.1), failure modes (Sect. 6.3.6.2), and dissipated energy (Sect. 6.3.6.3). Final remarks concerning the comparison between the efficiency of each strengthening solution are discussed in Sect. 6.3.6.4.

6.3.6.1

Out-of-Plane Force–Displacement Envelopes

Figure 6.24a and b show a comparison among the test results in terms of FOOP − dOOP envelopes of the specimens without prior damage (Inf_08, Inf_10 and Inf_13) and with previous IP damage (Inf_11 and Inf_12). In the same way, the relative secant stiffness evolution is shown in Fig. 6.24c and d. Additionally, Table 6.5 and Fig. 6.25 summarise the main results, as commented in the following remarks. Out-of-plane drift at control node (%) 1.74

3.48

5.22

Out-of-plane force FOOP (kN)

No previous IP drift

6.96

8.70

Out-of-plane drift at control node (%)

10.43 0.00 100

Inf_08 (As-built) Inf_10 (Strengthening solution A) Inf_13 (Strengthening solution B)

80

Cracking point Peak Load

60 Limit of the pneumatic atuators

Panel collapse

40 End of test

20

1.74

3.48

5.22

6.96

8.70

10.43

Inf_11 (As-built) Inf_12 (Strengthening solution B)

Previous IP drift: 0.3%

Out-of-plane force FOOP (kN)

0.00 100

Cracking point Peak Load

80

60 Limit of the pneumatic atuators

40

Limit of the pneumatic atuators

20

Inf_08

0 0

20

40

60

80

100

0

120

0

Out-of-plane displacement at control node dOOP (mm)

20

40

a)

100

120

Inf_11 (As-built) Inf_12 (Strengthening solution B) Previous IP drift: 0.3%

1.0

0.8

0.8

Relative stiffness

Relative stiffness

80

b) Inf_08 (As built) Inf_10 (Strengthening solution A) Inf_13 (Strengthening solution B) No previous IP drift

1.0

60

Out-of-plane displacement at control node dOOP (mm)

0.6

0.4

0.2

0.6

0.4

0.2

0.0

0.0 0

20

40

60

80

100

120

Out-of-plane displacement at control node dOOP (mm)

c)

0

20

40

60

80

100

120

Out-of-plane displacement at control node dOOP (mm)

d)

Fig. 6.24 Assessment of the efficiency of the strengthening techniques: a force displacement envelope curves (specimens without prior damage); b force displacement envelope curves (specimens with previous IP damage); c relative stiffness (specimens without prior damage); and d relative stiffness (specimens without prior damage)

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6 Experimental Assessment of Strengthening Solutions to Prevent …

Table 6.5 Assessment of the efficiency of the strengthening techniques: comparisons of the results Parameters

Reference

Strengthening solution A

Reference

Strengthening solution B

Strengthening solution B

Inf_08 (OOP)

Inf_10 (OOP)

Inf_11 (IP-OOP)

Inf_12 (IP-OOP)

Inf_13 (OOP)

FOOP,max (kN)

44.15

77.10

30.72

75

92.31

FOOP,crack (kN)

22.60

47.23

23.09

44.90

59.11

FOOP,ult (kN)

37.96

39.40

13.33

39.16

42.61

kOOP,sec,ini (kN/mm)

20.32

18.90

7.97

7.44

80.92

dOOP,max (%)

2.21

2.35

4.47

2.75

2.55

dOOP,crack (%)

0.11

0.33

0.67

0.71

0.17

dOOP,ult (%)

3.20

6.31

8.87

8.91

8.70

12

100 Strengthening solution B (OOP)

FOOP,crack

OOP strength FOOP (kN)

80

FOOP,ult

Strengthening solution A (OOP)

dOOP,crack Strengthening solution B (OOP) Strengthening Reference (IP-OOP) solution B (IP-OOP)

dOOP,max dOOP,ult

Strengthening solution B (IP-OOP)

9

OOP drift dOOP (%)

FOOP,max

60 Reference (OOP)

40 Reference (IP-OOP)

Strengthening solution A (OOP)

6

Reference (OOP)

3

20

0

0 Inf_08

Inf_10

Inf_11

Specimens

a)

Inf_12

Inf_13

Inf_08

Inf_10

Inf_11

Inf_12

Inf_13

Specimens

b)

Fig. 6.25 Assessment of the efficiency of the strengthening techniques: a OOP strength FOOP ; and b OOP drift dOOP

Initial secant stiffness The initial secant stiffness, kOOP,sec,ini , was calculated for each specimen, and from the results summarized in Table 6.5 it follows that the maximum value was reached by the strengthened panel Inf_13 with 80.92 kN/mm and the lowest value was achieved by the strengthened specimen Inf_12 with 7.44 kN/mm. The no expected result obtained by the panel Inf_12 can be justified by the fact that it was firstly damaged during the IP test, which reduced the initial stiffness of the panel. Actually, comparing with the reference specimen, Inf_11, the result is quite similar, which allows to conclude that the strengthening solution does not contribute to the panel initial stiffness. The strengthened panels subjected only to pure OOP load (Inf_10 and Inf_13) achieved initial secant stiffness values of 0.93 and 3.98 times the value reached by the reference specimen.

6.3 Experimental Assessment of TRM Solutions to Improve …

253

First cracking From the analysis of the results obtained by each specimen, it is possible to observe that, regarding the force corresponding to the development of the first cracking, the panel Inf_13 reached the maximum value of 59.11 kN, 62% higher than the result obtained by Inf_08 that achieved the lowest cracking strength. From the comparison between the specimen Inf_10 and the reference specimen Inf_08, it is possible to conclude that the strengthening solution contributed to the increase of the cracking strength about 2.08 times. By comparing the specimen Inf_10 with the other strengthened specimen subjected to a pure OOP test, specimen Inf_13, it was observed that the strengthening solution B reached a cracking strength about 1.31 times higher. Regarding the comparison of the specimen Inf_12 and the reference one Inf_11, it is possible to observe that the strengthening solution improved the cracking strength about 1.95 times. Finally, concerning the impact of the damage caused by the IP test on the efficiency of the strengthening solution B, the cracking strength of the specimen Inf_12 reduced about 24% compared with the result of the specimen Inf_13. Interesting results were obtained regarding the OOP drift for which occurred the first macro-cracking, namely the maximum value reached by Inf_12 with 0.71%, which is justified by the fact that the Inf_12 was first subjected to the IP test. The lowest value was achieved by 0.11% by specimen Inf_08. The comparison between the strengthened specimens and the respective reference specimens shows that the strengthening solutions contributed to the increase of the dOOP,crack in all the examples between 1.06 and 3 times. Maximum peak stage Concerning maximum peak load, very satisfactory results were obtained by all the strengthened panels as proved by the increase between 1.7 and 2.44 times. The largest FOOP,max was achieved by Inf_13 with a value equal to 92.31 kN, 2.44 times higher than the as-built reference specimen (Inf_08). The specimen Inf_10 maximum OOP strength was 77.10 kN, 1.74 times higher than Inf_08 and 16% lower than the strengthened panel Inf_13. Regarding the last strengthened specimen Inf_12, the maximum strength equal to 75kN was 2.44 times higher than Inf_11 and only 18% lower than the result obtained by Inf_13, which is proof that the efficiency of the strengthening solution was not significantly affected by the prior damage. From the analysis of Fig. 6.25a, it is possible to observe that the strengthening was efficient, since all the strengthened panels reached higher maximum peak loads. The same trends can be noticed regarding the OOP drift corresponding to each stage (Fig. 6.25b). The specimens without prior damage, reached the maximum peak load for dOOP,max values between 1.06–1.15 times higher. On the other hand, concerning the specimens with previous damage, the maximum peak load of the strengthened specimen occurred for a dOOP,max 39% lower. Ultimate stage The ultimate stage of all the specimens was studied and it correspond to the last instant of the tests (FOOP,ult and dOOP,ult ), which in one case (Inf_08) was due to the

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panel collapse, in another case (Inf_10) was due to the risk of collapse and consequent protection of the instrumentation and in the remaining specimens correspond to the limit of the pneumatic actuators. From the point of view of the collapse prevention, it becomes clear that both strengthening solutions were effective. The test of panel Inf_10 did not allow to take solid conclusions concerning this aspect, since the test has stopped for earlier OOP displacements. However, the results obtained by the panels Inf_12 and Inf_13 show that the use of a proper TRM strengthening solution can prevent the panel OOP collapse and its consequences. The results are particularly good since the residual strength of the panels Inf_12 and Inf_13 were 39.16 kN and 42.61 kN, respectively which allow to highlight the following conclusions: (i) the negligible effect of the IP damage in the efficiency of the mesh for the OOP displacement; (ii) the increase of the OOP deformation capacity for larger displacements without collapse; (iii) increase of the ultimate OOP strength. In the case of the comparison among the panel Inf_13 and Inf_08 the major difference is the OOP drift level achieved by the strengthened panel, which is about 2.72 times higher without collapse.

6.3.6.2

Failure Mode and Collapse Prevention

An additional comparison among the presented test results can be carried out in terms of observed collapse prevention. Regarding the as-built specimens, the panel Inf_08 presented at the end of the test an OOP drift of 3.20%, characterized by the OOP collapse of the panel. The cracking pattern of the panel immediately prior to the collapse is shown in Fig. 6.26a. The cracking pattern was characterized by a trilinear shape, with the detachment of the panel from the top beam. The other asbuilt specimen, and first subjected to prior IP damage, the panel Inf_11 reached a maximum OOP drift equal to 8.87%. Similarly to Inf_08, a trilinear crack pattern developed throughout the test (Fig. 6.26c), with detachment from the top beam. The collapse of the panel did not occur. Regarding the strengthened specimens subjected to pure OOP tests (Inf_10 and Inf_13), both behaved similarly with distributed cracking throughout the panel without sliding failure of the mesh or detachment of the mesh-frame anchorage. However, some differences can be noticed namely the fact that the cracking pattern of the specimen Inf_10 was more concentrated from the mid-height to the bottom of the panel, by contrast with the specimen Inf_13, where it is visible a uniform distribution of cracking over the panel, as shown in Fig. 6.26b and e respectively. Finally, it is important to remark the differences between the strengthened panel Inf_12 and the as-built panel Inf_11. The use of the TRM solution, Inf_12, allowed reaching a similar OOP drift, through ensuring that the panel is protected by the mesh which increased the OOP strength. In the strengthened specimen, the spread cracking over the entire panel was visible, in contrast to the trilinear crack reached by the specimen Inf_08. The cracking pattern of the specimens Inf_11 and Inf_12 are shown in Fig. 6.26c and d, respectively.

6.3 Experimental Assessment of TRM Solutions to Improve …

255

Fig. 6.26 Assessment of the efficiency of the strengthening techniques: cracking pattern a Inf_08; b Inf_10; c Inf_11; d Inf_12 and e Inf_13

6.3.6.3

Dissipated Energy

The energy dissipated in each individual loading half-cycle and the cumulative energy dissipation throughout the whole test history were calculated and plotted in Fig. 6.27. From the plots of the energy dissipated in each cycle, some differences can be pointed: (i) some peaks can be identified in the strengthened panels that corresponded to sudden large displacements reached by the panels when the crushing occurred in the bricks located in the top and bottom rows (visible in the force–displacement

6 Experimental Assessment of Strengthening Solutions to Prevent … 10 Energy dissipated for each cycle Cumulative energy disspated Specimen Inf_08 (As-Built)

9

No previous IP drift

8

3.0 7 2.5

6

2.0

5 4

1.5

3 1.0 2 0.5

1

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Energy dissipated for each cycle Cumulative energy disspated Specimen Inf_10 (Strengthening solution A)

9

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8

3.0 7 2.5

6

2.0

5 4

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3 1.0 2 0.5

1

0.0 0

5

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a)

Previous IP drift: 0.3%

8

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1

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55

Dissipated energy for each cycle (kN.m)

9

10

35

0 40

10

4.0

Cumulative energy dissipated (kN.m)

Dissipated energy for each cycle (kN.m)

Energy dissipated for each cycle Cumulative energy disspated Specimen Inf_11 (As-built)

5

30

b) 10

0

25

Number of cycles

Number of cycles 4.0 3.5

20

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3.5

Energy dissipated for each cycle Cumulative energy disspated Specimen Inf_12 (Strengthening solution B)

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0

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3.5

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256

0 55

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c)

d) Dissipated energy for each cycle (kN.m)

3.5

Energy dissipated for each cycle Cumulative energy disspated Specimen Inf_13 (Strengthening solution B)

9

No previous IP drift

8

3.0 7 2.5

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Cumulative energy dissipated (kN.m)

10

4.0

0 45

Number of cycles

e)

Fig. 6.27 Assessment of the efficiency of the strengthening techniques: dissipated energy per cycle: a Inf_08; b Inf_10; c Inf_11; d Inf_12 and e Inf_13

response); and (ii) the cycle with largest energy dissipation in the specimen Inf_08 was precisely the last one, when the panel reached sudden increase of the OOP displacement. The cumulative energy dissipation was computed not considering the peak of energy dissipation, which is visible in the response of the specimens Inf_10, Inf_12 and Inf_13. From that, the total cumulative energy dissipation reached by the specimen Inf_08, Inf_10, Inf_11, Inf_12 and Inf_13 were, respectively, 1.9, 6.2, 3.9, 8.0 and 9.2 kNm. The cumulative energy dissipation evolution is plotted in Fig. 6.28, from which it is possible to observe that, for the same OOP displacement demand, the panels Inf_10 and Inf_13 dissipated similar levels of energy. From the comparison between

6.3 Experimental Assessment of TRM Solutions to Improve … 8

Cumulative energy dissipation (kN.m)

Fig. 6.28 Assessment of the efficiency of the strengthening techniques: cumulative energy dissipation

257

Inf_08 (As-built; OOP test) Inf_10 (Strengthening solution A; OOP test) Inf_11 (As-built; IP-OOP test) Inf_12 (Strengthening solution B; IP-OOP test) Inf_13 (Strengthening solution B; OOP test)

7 6 5 4 3 2 1 0 0

20

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OOP displacement at control node (mm)

the panels without previous IP damage, it can be observed that for an OOP displacement around 30 mm, the specimens Inf_10 and Inf_13 reached a cumulative energy dissipation 45 and 48% higher than that of Inf_08, respectively. Regarding the total cumulative energy dissipation, it is clear the efficiency of the strengthening, since it was reached by the strengthened specimens’ values 3.26 and 4.84 times higher, respectively Inf_10 and Inf_12. It is also visible that the strengthening solution B was more efficient, since the cumulative energy dissipation reached by Inf_13 reached 1.48 times higher than that of Inf_10. From the comparison of the results of the panels with previous IP damage, it is visible that the strengthening was efficient and contributed to the increase of the energy dissipation about 2.05 times.

6.3.6.4

Analysis of the Efficiency of the Strengthening Solutions

To conclude the analysis of the efficiency of the strengthening solution it is important to explore and discuss the force–displacement curves, where it is possible to verify the effect of the reinforcement in the response parameters, such as initial stiffness, cracking point (dOOP,crack and FOOP,crack ), maximum peak load (dOOP,max and FOOP,max ) and ultimate point (dOOP,ult and FOOP,ult ). Table 6.6 presents a summary of strengthening factors (ratio between the as-built specimens and the strengthened ones or ratio between the strengthened specimens). Starting from the analysis of the results obtained by the panels without previous IP damage (As-built: Inf_08; strengthened panels: Inf_10 and Inf_13) that are plotted in Fig. 6.29a–c, the following observations can be drawn: . From the comparison between the specimen Inf_08 and Inf_10 force-displacement curves, shown in Fig. 6.29a, it is possible to observe the strengthening solutions did not contributed to the increase of the initial stiffness. In fact, the initial stiffness of the panel Inf_10 was 7% lower. The first crack occurred for in the strengthened

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Table 6.6 Assessment of the efficiency of the strengthening techniques Parameters

Ratio Inf_10/Inf_08

Ratio Inf_13/Inf_08

Ratio Inf_10/Inf_13

Ratio Inf_12/Inf_11

Ratio Inf_12/Inf_13

FOOP,max

1.75

2.09

0.84

2.44

0.81

FOOP,crack

2.10

2.62

0.80

1.94

0.76

FOOP,ult

1.04

1.12

0.92

2.94

0.92

kOOP,sec,ini

0.93

3.98

0.23

0.94

0.09

dOOP,max

1.06

1.15

0.92

0.62

1.08

dOOP,crack

3.00

1.55

1.94

1.06

4.18

dOOP,ult

1.97

2.72

0.73

1.01

1.02

panel for an OOP displacement (dOOP,crack ) 3 times higher and for a corresponding load (FOOP,crack ) 2.10 times higher. The large efficiency of the strengthening is visible by the improvement of the maximum peak load (FOOP,max ) about 1.75 times that occurred for a similar OOP displacement. It is possible to observe that, despite the significant strength reduction after the peak load, the residual OOP strength is quite significant (39.4 kN) and the strengthening allowed to reach OOP displacement 1.97 (dOOP,ult ) times higher than those reached by Inf_08 without collapsing; . Concerning the comparison between the strengthened panel Inf_13 and the asbuilt specimen Inf_08, shown Fig. 6.29b, it can be observed that the initial stiffness of the strengthened one was 3.98 times higher. The first crack occurred in the panel Inf_13 for an OOP displacement (dOOP,crack ) 1.55 times higher and for a corresponding load (FOOP,crack ) 2.62 times higher than that of Inf_08. The strengthening solutions increased the panel maximum peak load in 2.09 times, which occurred for a dOOP,max 15% higher. Again, the strengthening allowed to reach a dOOP,ult 2.72 times higher with a residual load 1.12 times higher. In the response of the strengthened specimen, after the peak load it occurred a sudden increase of the OOP displacement with a reduction of the OOP load, however the presence strengthening prevented the possibility of collapse; . Regarding the assessment of the most efficient strengthening solution, the responses of the panels Inf_10 (Strengthening solution A) and Inf_13 (Strengthening solution B), are shown in Fig. 6.29c. In all the response parameters (except dOOP,crack ), the specimen Inf_13 is more efficient by achieving higher results than that of Inf_10. For example, the initial stiffness of the panel Inf_13 was 75% higher. Concerning the first crack occurrence, it occurred for and OOP displacement 8% higher and for a corresponding load 20% higher. The maximum peak load reached by the panel Inf_13 was 16% higher and occurred for an OOP displacement 8%. Smoother after-peak strength degradation was visible in the panel Inf_13 response. The ultimate load of the panel Inf_13 was 8% higher and occurred for a dOOP,ult 27% larger.

6.3 Experimental Assessment of TRM Solutions to Improve … Out-of-plane drift at control node (%)

Out-of-plane drift at control node (%) 1.74

3.48

5.22

6.96

8.70

10.43

0.00 100

Specimen Inf_08 (As-built) Specimen Inf_10 (Strengthening Solution A) No previous In-Plane drift Cracking point Peak Load Displacements not recorder Ultimate Load (out of range)

80

Out-of-plane force FOOP (kN)

Out-of-plane force FOOP (kN)

0.00 100

259

60

40

20

0

1.74

3.48

6.96

8.70

10.43

Specimen Inf_08 (As-built) Specimen Inf_13 (Strengthening Solution B) No previous In-Plane drift Cracking point Peak Load Ultimate Load

80

60

40

20

0 0

20

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120

0

20

40

a) 1.74

3.48

100

120

b) 5.22

6.96

8.70

10.43

0.00 100

Out-of-plane force FOOP (kN)

Specimen Inf_10 (Strengthening Solution A) Specimen Inf_13 (Strengthening Solution B) No previous In-Plane drift Cracking point Peak Load Displacements not recorder Ultimate Load (out of range)

80

80

Out-of-plane drift at control node (%)

Out-of-plane drift at control node (%) 0.00 100

60

Out-of-plane displacement at control node dOOP (mm)

Out-of-plane displacement at control node dOOP (mm)

Out-of-plane force FOOP (kN)

5.22

60

40

20

1.74

3.48

5.22

6.96

8.70

10.43

Specimen Inf_11 (As-built) Specimen Inf_12 (Strengthening Solution B) Previous In-Plane drift:0.3% Cracking point Peak Load Ultimate Load

80

60

40

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0

0 0

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120

Out-of-plane displacement at control node dOOP (mm)

20

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Out-of-plane displacement at control node dOOP (mm)

c)

d) Out-of-plane drift at control node (%)

Out-of-plane force FOOP (kN)

0.00 100

1.74

3.48

5.22

6.96

8.70

10.43

Specimen Inf_12 (Strengthening Solution B; 0.3% IP drift) Specimen Inf_13 (Strengthening Solution B; No prior IP drift) Cracking point Peak Load Ultimate Load

80

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40

20

0 0

20

40

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80

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Out-of-plane displacement at control node dOOP (mm)

e)

Fig. 6.29 Analysis of the strengthening techniques efficiency: force–displacement curves: a Inf_08 vs. Inf_10; b Inf_08 vs. Inf_13; c Inf_10 vs. Inf_13; d Inf_11 vs. Inf_12 and e Inf_12 vs. Inf_13

From the analysis of the force–displacement results of the panels subjected to previous IP damage (Inf_11 and Inf_12), shown in Fig. 6.29d, the strengthening increased all the strengthening parameters, namely FOOP,crack , FOOP,max , and FOOP,ult about 1.94, 2.44 and 2.94 times higher, respectively. From this comparison, it is visible that the strengthening is more effective for larger OOP displacements. Through the analysis of the initial stiffness, the strengthened panel reached a value 6% lower. It is also important to say that, concerning the discussion of the OOP

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6 Experimental Assessment of Strengthening Solutions to Prevent …

displacement parameters, the dOOP,crack and dOOP,ult were similar but the dOOP,max achieved by the specimen Inf_11 was 38% higher. To conclude, Fig. 6.29e shows the force–displacement curves of the specimens Inf_12 and Inf_13 curves, both strengthened with solution B. From this comparison, it is possible to assess the impact of prior damage in the strengthening performance. Through the observation of the force– displacement curves, it is possible to observe that the strength degradation due to the IP damage occurs for similar OOP displacement demands, which means that it corresponds to stiffness degradation. The cracking, maximum and ultimate strength reduced about 24%, 19% and 8%, respectively. To analyse the efficiency of the strengthening solution A for larger displacements, it is mandatory the realization of an additional test until reached the maximum OOP displacement as possible (only limited by the pneumatic actuators capacity).

6.4 Final Considerations The present chapter aimed to present an experimental assessment of TRM strengthening solutions to improve the OOP behaviour of infill walls. This chapter started with a testing campaign that was carry out to assess the performance of TRM solutions in small wallets subjected to flexural strength tests. The influence of using different types of mesh was assessed. Two types of flexural strength tests were carried out, namely four-point flexure strength tests and cantilever flexural strength tests. The second component of this chapter was the realization of a testing campaign of quasistatic OOP tests on full-scale infill panels aiming at evaluation the efficiency of two TRM strengthening solutions to prevent the OOP collapse. From the testing campaign of the strengthened infill masonry wallets the following conclusions can be drawn: . The TRM strengthening technique revealed to be effective to improve the flexural capacity of the panels, both in terms of strength and deformation; . The selection of the mesh should be performed taking into account the type of masonry unit used to build the panel, since if weak brick units such as HCHB are used, it will not explore all of the tensile strength characteristics of the mesh. The bricks will crush before the rupture of the mesh. It was observed that the rupture of the mesh occurred only when it was used the weak PP mesh; . The dimension of the specimens, which were designed according to the code standard, affected the conclusions of the testing campaign since some specimens failure were due to shear effect. The geometric dimensions of the specimens should be revised new tests should be made to clarify the findings presented in this chapter; . Additional tests should be also performed by testing infill wallets strengthened with the mesh used in the strengthening solution B, presented in Sect. 6.3.2.2, and relates with results obtained in the OOP testing of the full-scale infill walls.

6.4 Final Considerations

261

The second goal of this chapter was the execution of a testing campaign to assess the efficiency of TRM strengthening solutions to prevent the collapse of infills OOP behaviour. An experimental campaign was carried out comprising five nominally identical full-scale specimens were built. Two as-built specimens were used as references, one with and one without prior damage caused by IP test. The remaining three specimens were strengthened using TRM technique. Two solutions were tested by varying the type of mesh (strong and medium) and the anchorage (using and not using steel plates to distribute the mesh-frame anchorage. All the tests have been performed by imposing a half-cyclic (loading–unloading-reloading) history of displacements in the out of plane (OOP) direction by means of a uniform distributed load provided by small pneumatic jacks. From the testing campaign the following conclusions can be drawn: . Both strengthening solutions revealed to be very efficient since they prevented the collapse of the infill panels. It was observed that the use of steel connectors for the mesh-frame anchorage was very effective since it was not observed any detachment or failure. The use of the steel plate improved the mesh performance by reducing the sliding failure of the mesh in the regions surrounding the steel connector; . The use of the strong mesh did not presented better results in terms of strength than the observed with the medium strength mesh. Although, this analysis cannot be directly extracted because different anchorage solutions were adopted, still due to the large strength characteristics and the lower ductility of the strong mesh A, it was not totally explored as proved by the cracking pattern of the specimen Inf_10; . The results of the strengthened specimens revealed that the TRM solutions improved about 1.75–2.44 times the maximum strength, 1.04–2.94 times the ultimate load, until 4 times higher initial stiffness, until 2.72 times the deformation capacity of the panel and improved the energy dissipation capacity at least 2 times, when compared with the as-built specimens tested in similar conditions. From the application process of the strengthening material and from the results herein presented, some aspects can be improved and tested in future specimens, namely: . The width dimension of the steel plate can be increased (around 1 cm) in order to better distribute the anchorage to the mesh; . Optimization of the number of metallic connectors and the dimension of the connector. It should be also take into account that the definition of the position of the metallic connectors must consider the RC cover of columns and beams to avoid their application along the steel bar alignment; . The definition of the plastic connectors for the panel-mesh anchorage should be done according to the bed-joints alignments, since the anchorage is more effective when the drill were made in the bed joints; . Optimization of the mesh characteristics. Since the major aspect is the anchorage of the mesh to the frame, the mesh can be adjusted to one with lower strength properties and more ductility.

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References 1. CEN (1999) EN 1052-2: Methods of test for masonry. Determination of flexural strength 2. De Risi M, Furtado A, Rodrigues H, Melo J, Verderame G, Arêde A, Varum H, Manfredi G (2020) Experimental analysis of strengthening solutions for the out-of-plane collapse of masonry infills in RC structures through textile reinforced mortars. Eng Struct (In Press) 3. CEN (2006) EN 196-2006 Methods of testing cement (European Committee for Standardization)

Chapter 7

Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

7.1 Introduction For the assessment of infilled RC frame structures, the non-linear behaviour induced by earthquakes should be considered [1, 61]. Different techniques are available in the literature to simulate the response of infilled frames, from simplified macromodels to refined detailed-models [1, 2]. In many cases, it is not suitable to adopt refined models for the non-linear analysis of complex structures when subjected to earthquakes. Thus, for the simulation of the response of infilled frame structures, considering the infill walls and their interaction with the surrounding frame elements, the adoption of simplified models is unavoidable. The present chapter aims first to present the development and implementation of a simplified macro-modelling approach to simulate the infill walls combined IP and OOP behaviour in the computer software OpenSees [3]. An element removal algorithm that allows removing elements during an earthquake simulation, accounting for interaction of both IP and OOP behaviour, was included in this modelling approach. The calibration of the numerical model to simulate the IP behaviour of panels, with and without openings, is presented within this section. The accuracy of the modelling approach will be assessed in terms of force–displacement and energy dissipation responses’ comparison between the numerical and experimental results from three different testing campaigns. The calibration methodology will be presented and discussed. The second part of this chapter presents a case study on the seismic vulnerability assessment of an eight-storey building that will demonstrate the expected effects of considering the presence of the infill panels. For this study, non-linear incremental dynamic analyses (IDA) were carried out in the software OpenSees. Three 3D models were built considering different modelling approaches, namely: (i) bare frame model (only infill gravity load is assumed), (ii) model with infill walls, by simulating only their IP behaviour, and (iii) model with consideration the infill walls by simulating

© The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 A. F. Furtado, Seismic Vulnerability Assessment and Retrofitting Strategies for Masonry Infilled Frame Building, Springer Theses, https://doi.org/10.1007/978-3-031-20372-5_7

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7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

their combined IP-OOP behaviour. Three different parametric studies were carried out to study the following topics: • Effect of the infill panels in the seismic vulnerability assessment; • Effect of the infill panels in the floor response spectra; • Effect of the infill panels in the mainshock-aftershock earthquake sequence. Starting from the first parametric study, the seismic vulnerability assessment was performed by the analysis of the IDA results, from which each maximum inter-storey drifts ratios (ISDMAX ) were extracted and compared with drift limits corresponding to the different damage states proposed by several literature proposals. The fragility curves for slight, light, moderate, extensive, partial collapse and collapse damage states of the structure are presented and discussed. In the second parametric study, as the title suggests, the effect of the infill panels in the floor seismic response spectra (FRS) of the building structure is analysed. For this purpose, non-linear IDA allowed extracting the floor accelerations’ and velocities’ time histories from which the peak floor accelerations (PFA) and peak floor velocities (PFV) were obtained. This study analysed the peak infills’ OOP accelerations (PIA) and velocities (PIV). The PFA and PIA in-elevation amplifications are presented, discussed, and compared with the results obtained through the Eurocode 8 [4] proposal to estimate the peak acceleration of the non-structural elements. The third and last parametric study is focused on methodology for assessing the seismic vulnerability of damaged RC structures that were first subjected to a mainshock event. A series of aftershocks with different magnitudes were applied and the role of the infill panels in the building structure response was evaluated.

7.2 Brief Overview on Numerical Modelling Approaches: From Simplified Macro-models to Detailed Modelling Different approaches are available in the literature to simulate the infill panels’ seismic behaviour, which can be divided into two groups: simplified macro models and detailed micro-modelling approaches. The first involves simplified models based on a physical understanding of the behaviour of infill panels. A small number of struts are used to represent the effect of the infill panels on the structural response of buildings when subjected to lateral loadings. This approach allow a satisfactory representation of the global behaviour of the panels and their interaction/influence with the RC elements. One of the major advantages is that it requires low computational effort. The second approach involves models in which the panel is discretized into numerous elements to consider the local effects in detail, allowing the representation of the infill panel at the local level. Parameters such as cracking pattern, collapse mechanism and ultimate load are well captured. This type of modelling approach is useful to calibrate the global models. Nonetheless, this strategy requires high computational and time efforts.

7.2 Brief Overview on Numerical Modelling Approaches: From Simplified …

265

Throughout the next sub-sections, a brief state-of-art review is presented and discussed concerning modelling strategies to simulate de infill panels’ seismic behaviour and the recent advances in terms of both modelling methodologies.

7.2.1 Simplified Macro-models The macro-modelling with equivalent diagonal struts was originally developed to capacitate numerical analysis models of infilled frames with higher shear stiffness. From its evolution, with multi-strut models it was possible to integrate shear and tensile stresses within the contact length between infill panel and frame. Models have started to become more complex, with some considering the reduction of stiffness and strength under dynamic loads, or other equivalent approaches to consider the shear slip at the middle of the infill walls. One of the aspects yet to be developed is the OOP behaviour itself, an even more important issue when combined with the diagonal cracking created by IP demands on the masonry infill walls. Polyakov [5] started with the proposal of an equivalent strut model to simulate the infill walls behaviour. The proposal was based on experimental observation studies on steel frames with focus on normal and shear stresses on the infill walls, in which it was said that the stresses were only transferred by the compression corners of infillframe interfaces from the structure to the non-structural elements. From that work, the authors developed a numerical technique to estimate the load intensity to create diagonal cracking. Holmes [6] improved the previous concept being the first author to propose a formulation for the diagonal strut. The proposed formula to calculate the equivalent strut width is a simplified approach, calibrated for steel frames with brickwork and concrete infill walls. It triggered several other studies to define the width more accurately. This simplified model considered deformation and ultimate strength of the global infill panel. Later, [7] compared the results from the experimental work developed in steel frame infilled specimens and concluded that the equivalent strut width depends on the infill-frame contact. The author observed that a larger contact length results in the increase of stiffness. From that, the author proposed different widths for the equivalent strut, which depend of the geometry of the panel. Later some studies were developed with the main purpose of calibrating the equivalent strut width and lateral stiffness of infilled frames [8, 9, 62, 63]. The consideration of the openings on infill panels was first considered by [10] that proposed an equation to reduce the stiffness provided by the infills. Klingner et al. [64] first introduced the double-strut model concept by the simulation of an infill panel with two equivalent struts. Besides the stiffness and strength capabilities, this proposal was also innovative since it was one of the first works to incorporate the hysteretic response on the elements for simulation of the cyclic behaviour and to include non-linear behaviour on a strut model. Additionally, the authors included a new proposal for the equivalent strut width. Leuchars et al. [65] proposed a modelling approach able to represent the response of panels with potential shear sliding failure. During the subsequent years, different authors [11, 12, 66,

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7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

67] focused their studies in the computation of the equivalent strut width, stiffness and strength based on experimental data. Thiruvengadam [13] introduced the multi-strut concept, to overcome some of the limitation regarding the inaccuracy of the single-strut models to compute bending and shear forces in RC elements. The author did not propose a limit number to the struts and their configuration, the only criteria was the capability to simulating the effect of the openings. Syrmakezis et al. [68] proposed a five parallel strut model in each direction to study both stiffness and strength of infilled RC frames. Chrysostomou [14] proposed a model comprising three parallel struts in each direction. The length considered for the connection of the struts to the RC elements was assumed equal to their plastic hinge length. This model predicted also the strength and stiffness degradation. More recently, Hamburguer et al. [69] proposed a multi-strut model accounting for openings. In 1995, Saneinejad et al. [70] proposed a numerical model that consisted on two diagonal and parallel struts in each direction of the infill wall, distributing the loads to the column elements. One of the characteristics of the model is the ability to consider the infilled frame ductility limitation, the frame aspect ratio, the shear stresses on infill-frame interfaces and the relative strengths between column and beam elements. One of the most well-known numerical models is the one proposed by Crisafulli et al. [71]. It presented an integration of struts and springs to compute two phenomena independently, namely: (i) diagonal cracking and corner crushing; and (ii) shear sliding. The model considers six strut members using hysteresis rules. It consists of two diagonal and parallel struts in each direction, which carry the axial loads on the panel, and another pair to simulate shear from the top and bottom of the panel, which are activated in each direction, depending on the activation due to axial compressive loads while the panel is deformed. Additionally, [15] compared different one-strut, double-strut and triple-strut models, concluding that the doublestrut model was the most balanced of the strategies, achieving accurate results without too much complexity in terms of calibration and computational efforts. According to the authors, the model finds its limitation in the connection to beam-column joints that does not provide accurate development of bending moments and shear forces on the structural elements neither the consideration of the panel OOP behaviour. This numerical model was later implemented in the software SeismoStruct [16] by [17]. In this implementation, the compressive struts distribute the loading through internal and dummy nodes, located on the joints and the columns, considering the actual contact of the infill panel when it is deformed. Few years later, the Crisafulli model was also implemented in OpenSees [3] by [18] in which the authors simulated the experimental behaviour of the IP behaviour of infilled RC frames. The authors also performed a parametric study to evaluate the sensitivity of the numerical results to the selection of the modelling parameters (in particular the empirical ones). El-Dakhakhni [19] proposed a model with three diagonal struts on each direction, one in the diagonal of the panel and the other two non-parallel in the off-diagonal. According to the researchers, this approach was better suited to compute the wall stiffness and to describe the development of stresses along the frame elements when

7.2 Brief Overview on Numerical Modelling Approaches: From Simplified …

267

compared to other models with less diagonal struts. The model was later implemented in ANSYS software [20]. The frame was modelled with elastic elements with the non-linearities lumped on the frame joints with springs. This simplified non-linear model was capable of computing the frame-infill interaction and corner crushing failure mechanism. Regarding the simplified macro-models which simulate the IP behaviour of infill panels, [2] presented a literature review. Until 2009, the simulation of the infill panels was only performed by considering their IP behaviour without any consideration of the OOP response. Kadysiewski et al. [72] proposed the first simplified model capable of simulating both IP and OOP behaviour of the infill walls, with a single diagonal beam-column element with a node at the mid-span having a concentrated mass to trigger the OOP inertia forces. Another novelty was developing and implementing an element removal algorithm, which allowed removing the panel during an ongoing analysis when the midpoint displacement exceeded a displacement curve interaction. The IP-OOP displacement interaction curve suggested by the authors is a 3/2-power curve. Furtado et al. [21] proposed an equivalent double-strut model comprising four diagonal struts and one central element where which simulates the combined IP and OOP behaviour. For this, a non-linear IP uniaxial material model was assumed for the central element. The mass of the panel was distributed into the two central nodes. It was assumed by the authors a linear elastic OOP behaviour. The model was implemented in the software OpenSees [3]. More details concerning this numerical model are provided in Sect. 7.3. Later, [22] also proposed a new macro element model for the simulation of the IPOOP response of infilled frames subjected to seismic actions. The model consists of two diagonals plus one horizontal and one vertical struts. Two fibre-section modelling beam-column elements represent each strut. The model can capture the arching action of the wall under an OOP load and the interaction between the IP and OOP actions. Ricci et al. [23] recently proposed a numerical model where each infill panel is simulated by a diagonal element representing their IP behaviour. Each of these elements is connected through a pinned joint to the surrounding frame and are provided with a central node that is connected to a second central node in which the mass participating in the first OOP vibration mode of the infill is lumped. The OOP backbone curves for low and medium–high IP drift levels. The main aim of the numerical model is to: (i) simulate the infills OOP behaviour, (ii) consider the OOP strength degradation due to IP damage and vice-versa; (iii) consider the OOP stiffness degradation due to IP damage and vice-versa, in order to correctly evaluate the acceleration and displacement demands on the infills during structural analyses. Mazza [24] proposed a five-element macro-model, with four support pin jointed (diagonal) truss elements with rigid axial behaviour, with inclination θ relative to the horizontal direction, and one (horizontal) truss element representing the hysteretic behaviour in terms of tensile and compressive axial forces. The central element is linked to the diagonal ones by two cylindrical hinges allowing the IP rotations, while spherical hinges connect the diagonal elements to the frame joints. The authors assumed that the axial behaviour of the five-element model reproduces the non-linear

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7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

IP behaviour, neglecting the in-plane flexural and shear stiffness. A distributed mass is also included in the floor mass of the beam supporting the infill panel. Concerning the OOP modelling, cylindrical hinges are used for restraining the OOP rotations of the four diagonal beams in the connection points with the central element, while the OOP rotations of these diagonal elements are allowed in the connection joints with the frame where spherical hinges are placed. Thus, the OOP non-linear response is reproduced in this model by using flexural and shear OOP behaviour of the fiveelement model. Two masses are applied in the two central nodes of the model. A bilinear hysteretic curve is proposed to simulate the OOP behaviour. Al Hanoun et al. [25] proposed a system with four elastic beam elements pinned to the joints of the RC frame elements and linked with a nonlinear axial link element. The numerical model considers the IP and OOP behaviour and their interaction in both directions. The OOP mass and stiffness were calculated according to the Kadysiewski et al. [72] proposal. The authors considered an OOP linear elastic behaviour and included the IP-OOP interaction curve through a linear displacement limit curve. The authors implemented and tested this numerical model in the software SAP2000 [26].

7.2.2 Detailed Micro-modelling Approach The detailed micro-modelling is a refined strategy in which all the elements composing the wall are modelled, namely masonry units and mortar joints, as volumetric elements and boundary or interface link-models to simulate the contact and friction conditions between the individual elements and frame. A simplified approach within this micro-modelling strategy may include reducing the number of elements by combining a brick with the surrounding mortar, which is connected to the rest by link-models. These approaches are expensive both on the modelling phase and computational demands, especially when applied to dynamic and non-linear analysis. On the other hand, the detailed modelling allows obtaining results that help to understand the behaviour at local level and the panel cracking pattern, which can be very useful for calibrating global models and performing parametric studies. This is an important advantage of the micro-models when compared with the simplified macro-models. With this modelling procedure the influence of each parameter in the infill panel seismic response can be assessed and quantified [1]. Micro-modelling approach started in 1967 with the work carried out by Mallick et al. [73], in which it was simulated the IP behaviour of an infilled RC frame, with particular focus in the frame-panel interface. The authors’ strategy was based on the modelling of the wall by rectangular elastic elements with two degrees of freedom per node. The frame-wall interaction was provided by the consideration of frictional shear forces to simulate slippage. A different approach was proposed by different authors such as [27, 28, 74, 75] with the introduction of the continuous-interface models’ concept, which basically can be applied to bed-joints by accounting for the interaction between the tangential

7.3 Development of a Simplified Modelling Approach to Simulate Infill …

269

and normal stress. Later, [28] proposed a model in which the Coulomb friction rule, tension cut off and compression strength are combined. From this model, the obtained damages are concentrated in the infill wall bed joints and in the middle of the masonry units. One of the simplifications proposed by the author, is the simulation of the infill panel as a three-phase material in which the units/mortar and their interfaces are modelled as continuous and discontinuous elements, respectively. For this purpose, the assumption made by the author was to use a simplified modelling for two-phase material, where the units are simulated by continuous elements, but the mortar and interfaces are lumped to discontinuous elements. Mohyeddin et al. [29] developed a generic three-dimensional discrete-finiteelement model that has been constructed for infilled RC frames using a commercial software to assess the in-plane and out-of-plane behaviour interaction. From the results, the authors found some differences between the behaviour predicted by the finite element model and the experimental results. The reasons behind these differences were justified by the authors as the combination of large coefficients of variation of masonry material properties and of the existence of weaker areas within the infill panel which were attributed to workmanship and that cannot be modelled. Finally, a more simplified approach was prepared by assuming one-phase materials, in which units, mortars and interfaces are combined into a continuum and homogeneous element. Chen et al. [76] developed a finite element model to simulate the IP behaviour of concrete masonry infills bounded by steel frames with openings. The authors proved that the model had the capability to simulate the experimental tests with high accuracy. Several other studies and efforts were carried out by other authors [30–32, 75, 77–82]. The literature review herein presented is just a brief summary of the advances regarding this modelling strategy since the main objective of the present chapter by contrast, the simulation of infill panels’ behaviour using a strut modelling approach. More information can be found in [1], which presents an extensive and indepth review on the state-of-the-art concerning the infill masonry micro-modelling approaches.

7.3 Development of a Simplified Modelling Approach to Simulate Infill Walls in OpenSees 7.3.1 General Considerations Considering field observations, it is desirable that numerical models constitute a fundamental tool to understand the infill walls behaviour when subjected to horizontal cyclic loadings. The accuracy of the simplified models is important, insofar as a correct representation of those elements may allow to achieve a more realistic seismic

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7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

vulnerability assessment of existing buildings (and the design of new ones) to prevent disasters through the adoption of strengthening solutions. This sub-section aims to present a simplified macro-model developed to be used in the software framework OpenSees [3] and consists of an upgrade of the [33] proposal which is an improvement of the commonly used equivalent bi-diagonal strut model. In this model, each masonry infill wall is simulated by four diagonal struts with rigid behaviour, and a central element where the non-linearity hysteresis is concentrated, with the two central nodes having the panel mass lumped in. The numerical model is defined with the elements and materials available in the OpenSees library [3], namely: the four diagonals are simulated through four elastic BeamColumns and one non-linear BeamColumn as the central element. This numerical model has the capacity to simulate the infill panels’ behaviour when subjected to combined IP and OOP cyclic loading. These two components are modelled independently although, when subjected to such biaxial cyclic action, they interact through an element removal algorithm that is implemented, as described in the next sub-section (Fig. 7.1). One of the most popular macro-models is the one proposed by [34], which was previously described in Sect. 7.2.1. The major differences between this model and the [34] model are: (i) the location of the diagonal struts, for which the Crisafulli’ proposal allows the transmission of the shear forces to RC columns, whereas the present one only transmits for the column–beam joint; (ii) the Crisafulli’ proposal reproduces better the infill walls’ shear behaviour through the pair of shear springs, however the presented model was calibrated also calibrated with IP experimental tests that revealed to capture very satisfactorily different types of infill wall nonlinear behaviour (failure mechanism characterized by diagonal cracking or by corner

Force-Based Beam-column element

RC beam

Elastic beam element

Force-Based Beam-column element Central node (mnode=0.405m panel)

Non-linear uniaxial element Infill panel IP hysteretic behaviour

RC beam

Force-Based Beam-column element

Fig. 7.1 Macro-model proposed to simulate the infill panels’ seismic behaviour

RC column

RC column

Central node (mnode=0.405m panel)

7.3 Development of a Simplified Modelling Approach to Simulate Infill …

271

crushing) as detailed in [35]; (iii) the arrangement of the diagonal struts and the central element of the presented macro-model can work as an entire block to simulate the panel OOP behaviour, and at the same time simulating the IP behaviour with good accuracy. The interaction between IP-OOP behaviour is an important difference between the present model and the Crisafulli model, as the former represents more realistically the expected behaviour of the IM when subjected to cyclic biaxial loadings. Recently, Ahmed et al. [83] validated the accuracy of the model presented in this chapter to simulate the IP behaviour of infilled RC frames. The authors carried out the definition of the numerical model according to the recommendations presented in the next sub-sections. The following sub-section presents the IP and OOP modelling methodology and assumptions.

7.3.2 Modelling of the IP Behaviour The IP behaviour is simulated through the central element with non-linear axial behaviour, which is characterized by a multi-linear curve representing the panel IP behaviour, defined by eight parameters, shown in Fig. 7.2, which characterize the following stages: (a) cracking (cracking force fi,c and drift dfi,c ); (b) intermediate “yielding” (yielding force fi,y and drift dfi,y ); (c) maximum strength, corresponding to the onset of crushing (fi,max and corresponding drift dfi,max ); (d) residual strength (fi,u ) and corresponding drift (dfi,u ). The hysteretic rules calibrated for infill models are controlled by three additional parameters: stiffness degradation α, pinching effect β and strength degradation γ. Within the OpenSees library, the pinching 04 uniaxial material model [3] was selected to be used to simulate the IP hysteretic behaviour and was assigned to the central element. This model is used to convey a material that represents a ’pinched’ fi,max fi,y

IM wall in-plane strength (MPa)

Fig. 7.2 Hysteretic material behaviour of the central element

fi,c

fi,u

df

i,c

df

i,y

df

i,max

Inter-storey drift (%)

df

i,u

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7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

load–deformation response and exhibits degradation under IP cyclic loading. Cyclic strength and stiffness degradation occurs according to three ways, namely: unloading stiffness degradation, reloading stiffness degradation, strength degradation. According to the above mentioned, this simplified macro-model does not account for short-column effects. Thus, for infilled frames where the short-column effect can be induced, a multiple strut model strategy should be adopted. After a careful review of previous studies carried out by different authors [36, 37, 84, 85], international codes’ recommendations and the numerical studies, it was found that the following values can be used to define the hysteretic curve for the infill panel behaviour: – The ratio between cracking and maximum strength (fi,c /fi,max ) is adopted as 0.55 and the cracking drift dfi,c is taken between 0.075 and 0.12% depending on the mechanical properties of bricks and mortar; – The intermediate-yielding force (fi,y ) and drift (dfi,y ) are determined as an intermediate point between the cracking displacement (fi,c , dfi,c ) and the maximum coordinate (fi,max , dfi,max ). The intermediate-yielding force (fi,y ) is taken between 65 and 75% of the maximum force (fi,max ), and the drift is taken between 0.15 and 0.35%; – The maximum strength (fi,max ) occurs for drift values approximately between 0.25 and 0.5%. The maximum strength can be calculated through Eqs. 7.1 and 7.2, where Ss is the infill masonry diagonal tensile stress obtained based on experimental tests, while t, Linf and hinf are the panel thickness, length and height, respectively. Both parameters were defined according to [38, 84],

f i,max

L in f × t × Ss = 0.818 (1 + CI C I = 1.925

L in f h in f

/ C I 2 + 1)

(7.1) (7.2)

– The post-peak strength degradation is based on the Dolsek et al. [85] recommendations estimating that the displacement at residual strength (dfi,u ) is five times the displacement of maximum strength, while the residual strength value (fi,u ) is about 20% of the fi,max . As future works, parametric studies should be carried out to calibrate these input parameters to define the hysteretic curve for panels built with different types of masonry units and to reduce the uncertainty and variability related to the procedure of defining the pinching curve. Finally, to complete the formulation of the pinching04 uniaxial material model it is also necessary to define parameters regarding the unloading stiffness degradation,

7.3 Development of a Simplified Modelling Approach to Simulate Infill … Table 7.1 Infill wall modelling: pinching4 parameters (unloading stiffness, reloading stiffness, strength degradation, pinching and energy degradation)

Unloading stiffness degradation

Reloading stiffness degradation

Strength degradation

Pinching

Energy degradation (gE)

273

αK1

0.70–0.80

αK2

0.65–0.70

αK3

0.60–0.70

αK4

0.65–0.70

αKlim

0.70–0.80

βD1

0

βD2

0

βD3

0

βD4

0

βDlim

0–0.2

ϒD1

1

ϒD2

0

ϒD3

1

ϒD4

1

ϒDlim

0–0.15

rDispP

0.7–1

fForceP

0–0.25

uForceP

0–0.1

rDispN

0.7–1

fForceN

0–0.25

uForceN

0–0.1

10

reloading stiffness degradation, strength degradation, pinching and energy degradation. Table 7.1 summarizes suggested values for the definition of the pinching4 parameters based in the guidelines provided by OpenSees Manual [3]. It is important, also, to clarify that this numerical modelling approach is not designed to simulate a specific failure mode, since the infill panel failure mode is unpredictable due to the large variability of the associated parameters, earthquake characteristics material uncertainty and workmanship. Nonetheless, the parameters herein proposed for the IP hysteretic behaviour of the central element were calibrated with experimental tests for which the predominant failure mode was diagonal cracking, as described later in Sect. 7.4. This IP numerical modelling approach can be adapted for each failure mode as required, since the non-linearity of the central element can be adjusted and be more representative of one specific type of failure mode. For that purpose, the experimental data can be used to calibrate the pinching properties of the central element and thus adapt the modelling approach for a specific goal.

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7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

7.3.3 Modelling of the OOP Behaviour The numerical modelling of the infills seismic response involves the need of simulation of the combined IP and OOP behaviour interaction which turns this process complex and hard to perform because both are dependent of several parameters such as: boundary conditions, mechanical and material properties of the composite brick–mortar, panel geometry and panel slenderness. As mentioned in Chap. 2, postearthquake survey assessments reported a significant number of OOP collapses, proving that this mechanism should be considered during the seismic safety assessment of existing buildings because it can introduce vertical or/and plan irregularities. Full importance should be given to the IP and OOP behaviour interaction, namely the influence of previous damage due to IP loading demands in the OOP capacity of the panel. The specific modelling of the OOP response through simplified macro-models is a difficult task, since there is not enough experimental data concerning this behaviour (considering all the relevant parameters) to support an accurate calibration of numerical models and, particularly, concerning to combined IP-OOP tests. As reported in the Chap. 2, the testing campaigns available in the literature involve a large number of variables, such as material variability and workmanship that increase the complexity of this process. The major concern regarding the inclusion of the OOP behaviour simulation in this modelling approach was to present a proposal based on simplified assumptions that could be calibrated in the near future based on the experimental tests. Due to that, some simplifications were assumed for the implementation of the OOP behaviour simulation in the present simplified macro-model, namely: – A linear OOP elastic behaviour is assumed for the central element; – Part of the panel mass is distributed to the central nodes of the numerical model. This part can be calculated as 0.81Mpanel , where Mpanel is the total mass of the infill panel and is divided by two (equal value for each node of 0.405Mpanel ); – It is assumed that the model has the same natural period as the original infill wall; for this, the OOP mass and bending stiffness values are calculated according to the [39] and ASCE-41 [40] recommendations, as well as the recommendations provided by Mosalam et al. [86]. In that way, the infill panel OOP inertia, associated with the panel bending stiffness, can be determined by Eqs. 7.3 and 7.4. The natural vibration period of the numerical model was calibrated with experimental values collected from the ambient vibration tests presented in Chap. 4, ( Ieq = 1.644 Iin f =

L diag h in f

)3 × Iinf

3 tin f × L in f

12

(7.3)

(7.4)

7.3 Development of a Simplified Modelling Approach to Simulate Infill …

275

In order to achieve a more realistic simulation of the infill panel’ seismic behaviour, an element removal algorithm developed by Kadysiewski et al. [72] was included. This algorithm removes automatically the collapsed elements during an ongoing simulation. For example, if the panel exceeds the IP-OOP drift limit surface (defined by the user) it is assumed that the panel has collapsed and, consequently, the algorithm removes automatically the panel mass and stiffness from the structure. With this modelling approach it is possible to obtain the panel OOP displacement, velocity and acceleration variation during a seismic event. The framework to implement the infill wall modelling in OpenSees and the operation mode of the element removal algorithm, are illustrated in Fig. 7.3. As a first assumption, it is assumed that the IP and OOP displacement interaction envelope follows a linear interaction, as can be observed in Fig. 7.4; the drift limits can be selected based on previous experimental tests, codes recommendations, depending on the user preferences. However, further experimental investigations should be performed to better characterize this interaction boundary and to define each direction drift limits according to the panel simulated.

Fig. 7.3 Simplified numerical model: framework layout scheme

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7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour IP drift (%)

δΙP

δΟΟP

OOP drift (%)

δΙP

OOP displacement (m)

Displacement limits IM wall numerical displacement

δΟΟP IP displacement (m)

Fig. 7.4 Infill wall displacement interaction law envelope to define element removal activation

7.4 Numerical Modelling Calibration The model calibration started with modelling and calibration of the IP behaviour by simulating the experimental tests of infilled RC frames with and without openings, both full-scale and scaled specimens. The main purpose was to improve this modelling approach capacity to simulate the IP behaviour of panels without and with openings (windows and doors). The experimental campaigns selected were those carried out by [41, 42, 44]. The numerical results are compared with the experimental ones in terms of force–displacement curves and energy dissipation. Later, some considerations are presented regarding the modelling calibration of the OOP behaviour.

7.4.1 In-Plane Behaviour—Panels Without Openings 7.4.1.1

Experimental Campaign by Pires [41]

The first assessment of the numerical model capability to simulate the infill panels IP behaviour was made through the simulation of a cyclic IP test performed by [41], on a single-bay infilled RC, scaled 2:3. The geometric characteristics of the frame, cross sections’ dimensions and the reinforcement detailing of the columns and beams are presented in Fig. 7.5. The axial load was applied on the top of the columns (to simulate the dead load) and cyclic horizontal displacements were imposed. The RC elements were simulated with beamwithhinges elements [3] and calibrated with the results of materials properties.

7.4 Numerical Modelling Calibration

277

a)

b)

c)

Fig. 7.5 Single-story, single-bay infilled masonry RC frame tested by [41]: a frame geometry b top beam cross sections and c columns’ cross-section dimensions and detailing of RC elements (units in meters)

The geometric dimensions of the specimen tested are 4.5 m long, 2 m high with cross section equal to 0.15 × 0.15 m for columns and 0.15 × 0.20 m for beams, as shown in Fig. 7.5. The input modelling parameters were calibrated with the material test results provided by the authors. The reinforcement steel used in the construction of the frame had an average Young elastic modulus of 190GPa and yielding strength of 434 MPa. From the concrete compression strength tests, it was obtained a mean value of 25.3 MPa and Young elastic modulus of 30GPa. Regarding the infill masonry properties, it was obtained a diagonal tensile strength of 0.27 MPa and a shear modulus of 0.48GPa. More detailed information about the material characterization tests can be found in [41]. The modelling parameters used to define the central element hysteretic curve were calculated according to the methodology described in Sect. 7.3.2 and were slightly adjusted to achieve the best possible response of the numerical model. The uniaxial material model curve is plotted in Fig. 7.6, and the input parameters are summarized in Tables 7.2 and 7.3, in terms of stresses and drifts. The backbone curve presents the positive and negative parameters, since it is the central element that simulates the infill panel seismic behaviour. All the backbone parameters were calculated according to the proposal presented in Sect. 7.3.2.

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7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

Fig. 7.6 Experimental campaign by Pires [41]: uniaxial material model backbone curve

IM wall strength (MPa)

fi,max+

Pinching04

1.0

fi,y+

0.5

fi,c

+

fi,u+

0.0

-0.5

fi,c-

fi,ufi,y-

-1.0 -1.5

fi,max-

-1.0

-0.5

0.0

0.5

1.0

1.5

Drift (%)

Table 7.2 Infill wall input modelling parameters for [41] fi,c (MPa)

dfi,c (%)

fi,y (MPa)

dfi,y (%)

fi,max (MPa)

dfi,max (%)

fi,u (MPa)

dfi,u (%)

0.61

0.03

0.82

0.20

1.03

0.45

0.35

1.25

Table 7.3 Experimental campaign by [41]: Infill wall input modelling parameters: pinching4 parameters

Unloading stiffness degradation

Reloading stiffness degradation

Strength degradation

Pinching

Energy degradation (gE)

gK1

0.78

gK2

0.69

gK3

0.69

gK4

0.69

gKlim

0.79

gD1

0

gD2

0

gD3

0

gD4

0

gDlim

0.1

gD1

1

gD2

0

gD3

1

gD4

1

gDlim

0.1

rDispP

0.8

fForceP

0.1

uForceP

0.01

rDispN

0.8

fForceN

0.1

uForceN

0.01

10

279

150

150

100

100

50

50

Base Shear (kN)

Base Shear (kN)

7.4 Numerical Modelling Calibration

0

-50

-100

-150 -0.15

Experimental Numerical -0.10

-0.05

0.00

0.05

0.10

0

-50

-100

Experimental Numerical 0.15

-150 -0.15

Top displacement (m)

-0.10

-0.05

0.00

0.05

0.10

0.15

Top Displacement (m)

a)

b)

Cumulative Energy Dissipation (kN.m)

45 40 35 30 25 20 15 10

Experimental Numerical

5 0 0.00

0.05

0.10

0.15

Top displacement (m)

c)

Fig. 7.7 Calibration of the numerical model results: Experimental campaign by [41] a Base shear Top displacement, b base-shear-top displacement envelopes and c Cumulative energy dissipation

From the global results, a good agreement was found between the experimental response in terms of force–displacement (Fig. 7.7a, b) and energy dissipation (Fig. 7.7c), which evidences the ability of this modelling approach to simulate the global hysteretic response of infilled frames. The base shear-top displacement envelope of the numerical model is about 5% higher than the experimental response for the same top displacement values. The dissipated energy was determined and once again the general model was less than 5% different than the experimental response, which is quite acceptable since it is a simplified model that is simulating the entire infill panel and its surrounding frame elements. In general, the numerical results are in good agreement with experimental results.

7.4.1.2

Experimental Campaign by Calvi et al. [42]

An experimental campaign was performed by Calvi et al. [42] on full-scale specimens in order to assess the experimental response of traditional and slightly reinforced infill

280

7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

panels made with HCHB units, for different intensity level earthquakes. The overall dimensions of each sample were about 4.5 × 3.0 m, width and height, respectively, and were designed as the lowest part of a four storey building, designed according to Eurocode 2 [43] and Eurocode 8 [4]. The details of the geometry and the reinforcement detailing are presented in Fig. 7.8. The specimens “test 01” (bare frame), “test 2” and “test 06” (infilled RC frames) were considered for the present study. Table 7.4 includes the material properties (concrete, steel and masonry) obtained from the characterization tests and provided by the authors. The uniaxial material model curves’ input parameters used to simulate the infills’ IP behaviour are summarized in Tables 7.5 and 7.6. More information relative to the experimental tests can be found in Calvi et al. [42]. First, the capacity of the numerical model was studied to simulate the bare frame model (Test 01), plotted in Fig. 7.9. The simulation was performed without the experimental data (displacement time history), which explains the low graphic quality of the shear-displacement curves of Test 01, 02 and 06 (later presented in Fig. 7.10) A

B

A'

B'

0.25m

3.0m

0.25m 4.2m

0.2m 0.3m

0.3m 0.2m

a) AA'

BB' 11Ø16

8Ø16

0.30m

8Ø22

0.25m 0.70m

0.70m

b)

0.30m

c)

Fig. 7.8 Experimental campaign by Calvi et al. [42]: single-story single-bay infilled masonry RC frame: a global specimen dimensions; b top and bottom beams’ section and reinforcement detailing and c columns’ section and reinforcement detailing

7.4 Numerical Modelling Calibration

281

Table 7.4 Experimental campaign by Calvi et al. [42]: Input material properties adopted for concrete, steel and masonry Value

Material properties Concrete Compressive strength

fc

30.1 MPa

Elastic modulus

Ec

29.6GPa

Steel Yield stress

fy

565 MPa

Elastic modulus

Ey

200GPa

Masonry Compressive strength

fm

1.1 MPa

Elastic modulus

Em

1.87GPa

fmo

5.54 MPa

Mortar Compressive strength

Table 7.5 Experimental campaign by Calvi et al. [42]: Input modelling parameters adopted for the infill walls Specimen

fi,c (MPa)

dfi,c (%)

fi,y (MPa)

dfi,y (%)

fi,max (MPa)

dfi,max (%)

fi,u (MPa)

dfi,u (%)

Test_02

0.57

0.09

0.76

0.22

0.98

0.5

0.23

1.5

Test_06

0.54

0.11

0.8

0.24

0.98

0.6

0.23

1.5

Table 7.6 Experimental campaign by Calvi et al. [42]: Infill wall input: pinching4 parameters

Parameters

Test_02 Test_06

Unloading stiffness degradation gK1

0.78

Reloading stiffness degradation

Strength degradation

0.78

gK2

0.69

0.69

gK3

0.69

0.69

gK4

0.69

0.69

gKlim

0.79

0.79

gD1

0

0

gD2

0

0

gD3

0

0

gD4

0

0

gDlim

0.1

0.1

gD1

1

1

gD2

0

0

gD3

1

1

gD4

1

1

gDlim

0.1

0.1 (continued)

282

7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

Table 7.6 (continued)

Parameters Pinching

Test_02 Test_06 rDispP

0.8

0.8

fForceP

0.1

0.1

uForceP

0.01

0.01

rDispN

0.8

0.8

fForceN

0.1

0.1

uForceN 0.01 Energy degradation (gE)

0.01

10

Fig. 7.9 A Calvi et al. [42] specimens numerical modelling: Bare frame Test 01: Base Shear top displacement results

since numerical results were overlapped the experimental curves presented by the authors in Calvi et al. [42]. Due to that, it was not possible to compare the energy dissipation of the experimental tests with the numerical results and thus to assess the numerical model efficiency. For the numerical simulations, it was considered the target displacements indicated by the author, however it is possible to observe a small difference between the target and the real experimental IP displacements experienced by the infilled frame. A good representation of the global frame response is achieved; for the smallest and largest drifts the numerical response captured well the strength degradation observed in the experimental response. Also, from the comparison between both

7.4 Numerical Modelling Calibration

283

(a)

(b) Fig. 7.10 Calvi et al. [42] specimens numerical modelling: a Full infill test 02, and b test 03 Shear force top displacement

284

7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

responses, it is possible to observe that the initial stiffness is captured by the numerical model with a difference around 2%. Concerning the maximum strength, the numerical model captured very well in the negative loading phase, but in the positive loading phase it is visible that the model did not reached the experimental displacement, which was due to the displacement law adopted in the pre-processor stage. Regarding the difference between the experimental and the numerical maximum strength, a numerical result was 3% lower in the negative loading phase and a difference lower than 1% in the positive loading phase. It was impossible to compare the energy dissipation during the experimental test and the numerical model result. From the numerical simulation results of the infilled frame specimens, Test 02 and Test 06 as plotted in Fig. 7.10a, b, it can be observed that, regarding the initial stiffness both tests exhibit good approximations: in the Test 06 the a difference was about 2% and Test 02 it was around 5%. After the initial stage, there is a slight difference between the second loading stiffness of the experimental result and the numerical one. This difference is observed in both tests’ responses. Regarding the capacity of the numerical model to simulate the strength capacity of the infilled frames, different observations can be pointed out: (i) better results were obtained for Test 02, in terms of both maximum strength and each cycle positive and negative strength; (ii) the maximum strength in the positive loading phase of the Test 06, captured by the numerical model, was 8% lower; and (iii) the strength degradation was well captured in both tests, but better results were obtained for the Test 02.

7.4.2 In-Plane Behaviour—Panels with Openings 7.4.2.1

Experimental Campaign by Kakaletsis et al. [44]

The experimental tests performed by Kakaletsis et al. [44] were composed by 1/3 scale tests of one-storey, one-bay RC frame with HCHB units. The reinforced concrete frame represents a typical ductile concrete construction, built following the currently used code standards in Greece. The dimensions of the frame and cross sections and the reinforcing detailing are presented, in Fig. 7.11. The mean compressive strength of the concrete was 28.5 MPa and the yield stress of the longitudinal rebars was 390.4 MPa. In this experimental campaign, two types of bricks were used: weak and strong. For the present study, only the weak brick infill tests were simulated. The mortar compressive strength was about 1.5 MPa and the compressive strength perpendicular to the holes and the respective elastic modulus was 2.6 MPa and 660.7 MPa (Table 7.7). The experimental campaign was composed by seven specimens, but in the present study it were considered only four experimental tests. Beyond the bare frame tests (Test B), only the specimens with weak infill were studied, particularly some with different infill panel layouts: (a) Full infill (test S), (b) Infill with window (test

7.4 Numerical Modelling Calibration

B

A'

B'

BB'

AA'

00

A

285

200

0.20m

0.20m

800

7Ø5.6

0.80m

7Ø5.6 150 0.15m

150 0.15m

b) 500

0.50m

0.15m

0.25m 0.15m

1.20m

8Ø5.6

150 0.30m 0.15m

0.15m 0.25m

a)

c)

Fig. 7.11 Kakaletsis et al. [44] single-storey and single-bay infilled masonry RC frame: a frame geometry, b cross sections dimensions and detailing of top beam and c columns

Table 7.7 Experimental campaign by Kakaletsis et al. [44]: Numerical modelling parameters adopted for the materials properties

Value

Material properties Concrete Compressive strength Elastic modulus

fc Ec

28.5 MPa 29.4GPa

Steel Yield stress Elastic modulus

fy Ey

390.4 MPa 200GPa

Masonry Compressive strength Elastic modulus

fm Em

2.6 MPa 0.661GPa

Mortar Compressive strength

fmo

1.5 MPa

WO2), and (c) Infill with door (test DO2), as shown in Fig. 7.12. More information concerning the experimental tests can be found in Kakaletsis et al. [44]. Table 7.8 presents the values adopted for the numerical modelling of the infill panels. All the parameters were obtained according to the methodology described in [45]. Based on the values presented in Tables 7.8 and 7.9, for the infill specimens with openings, the strength reduction factor suggested by [45] is equal to the ratio between the opening area and the panel area. The corresponding uniaxial material model curves were adjusted to the experimental response to capture the real behaviour of the structure. The numerical base shear-top displacement response for the four studied cases are presented in Fig. 7.13 and compared with the corresponding experimental ones. Once again, it was not possible to compare the energy dissipation of the experimental tests with the numerical results and to assess the efficiency of the numerical model

7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour A

B

A'

B'

00

286

0.20m

800

0.80m

500

0.50m

0.25m 0.15m

0.15m 0.25m

1.20m a)

0.20m

0.13m 0.33m

0.33m

0.50m

0.25m 0.15m

0.30m

0.45m

0.45m

0.15m 0.25m

b)

Fig. 7.12 Experimental campaign by Kakaletsis et al. [44]: specimens dimensions a Test S—Full infill, b Test WO2—Infill with window, and c Test DO2 Infill with Door

7.4 Numerical Modelling Calibration

287

0.20m

0.16m

0.64m

0.50m

0.25m 0.15m

0.30m

0.45m

0.45m

0.15m 0.25m

c)

Fig. 7.12 (continued) Table 7.8 Experimental campaign by Kakaletsis et al. [44]: input modelling parameters adopted for the infill panels Specimen

fi,c (MPa)

dfi,c (%)

fi,y (MPa)

dfi,y (%)

fi,max (MPa)

dfi,max (%)

fi,u (MPa)

di,u (%)

S

1.09

0.08

1.43

0.22

1.88

0.49

0.47

1.5

WO2

0.77

0.09

1.02

0.23

1.50

0.49

0.47

1.5

DO2

0.83

0.085

0.96

0.25

1.41

0.50

0.47

1.5

Table 7.9 Experimental campaign by Kakaletsis et al. [44]: input modelling parameters concerning the pinching4 curve S

WO2

DO2

gK1

0.81

0.80

0.78

gK2

0.72

0.70

0.69

gK3

0.72

0.70

0.69

gK4

0.72

0.70

0.69

gKlim

0.80

0.80

0.79

gD1

0

0

0

gD2

0

0

0

gD3

0

0

0

Parameters Unloading stiffness degradation

Reloading stiffness degradation

(continued)

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7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

Table 7.9 (continued) Parameters gD4 Strength degradation

Pinching

Energy degradation (gE)

S

WO2

DO2

0

0

0

gDlim

0.1

0.1

0.1

gD1

1

1

1

gD2

0

0

0

gD3

1

1

1

gD4

1

1

1

gDlim

0.1

0.1

0.1

rDispP

0.8

0.8

0.8

fForceP

0.1

0.1

0.1

uForceP

0.01

0.01

0.01

rDispN

0.8

0.8

0.8

fForceN

0.1

0.1

0.1

uForceN

0.01

0.01

0.01

10

because the authors did not provided the experimental force–displacement results. For the numerical simulations, it was considered the target displacements indicated by the author, however it is possible to observe a small deviation between the target and the real experimental IP displacements experienced by the infilled frame. Globally, the numerical base shear-top displacement curves follow the experimental response of each specimen. Starting from the Test B (Fig. 7.13a), it is visible a good agreement between the numerical and experimental initial stiffness, although 12% difference is observed between the positive maximum strength, being the obtained in the numerical model the lowest one. Regarding the negative direction, the maximum strength is captured by the numerical model, but for larger topdisplacement. It is not visible in the numerical response any strength degradation, which slightly occurs in the experimental test after the peak load. Concerning the test S (Fig. 7.13b), the initial stiffness is slightly underestimated in the positive direction, around 5%, and well captured in the negative direction. The numerical and experimental maximum strength in the positive loading phase totally match, but in the negative loading phase it is overestimated by the numerical model, by around 12%. The strength degradation is slightly lower than the experimental response, although following a similar pattern. It is visible that for each displacement peak, the drop between the 1st and 2nd cycle maximum strength is lower in the numerical response. Finally, the Test S ultimate strength reached by the numerical model is 1.35 times higher in the positive loading phase and about the same value in the negative loading phase. Concerning the specimens with openings (Tests WO2 and DO2) it can be observed, that in both tests the numerical model overestimates the initial stiffness, in the negative loading phase, around 5%.

7.4 Numerical Modelling Calibration

289

90

Test B Exp Num

70 50

Base Shear (kN)

30

10 -40

-30

-20

-10

-10 0

10

20

30

40

20

30

40

-30

-50 -70 -90 Top Displacement (mm) a)

Test S

90

Exp Num

70 50

Base Shear (kN)

30 10 -40

-30

-20

-10

-10 0

10

-30 -50 -70

-90 Top displacement (mm) b)

Fig. 7.13 Kakaletsis et al. [44] specimens numerical modelling results: a Test 01—Bare Frame b Test 02—Full infill c Test 03—Infill with window d Test 04 Infill with Door

290

7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour 90

Test WO2 Exp Num

70 50

Base Shear (kN)

30 10

-40

-30

-20

-10

-10 0

10

20

30

40

-30 -50 -70

-90

Top Displacement (mm) c) 90

Test DO2 Exp

70

Num 50

Base Shear (kN)

30 10

-40

-30

-20

-10

-10 0

10

-30 -50

-70 -90

Top displcament (mm) d)

Fig. 7.13 (continued)

20

30

40

7.4 Numerical Modelling Calibration

291

Looking at the Test WO2 (Fig. 7.13c), the numerical response slightly overestimated the maximum strength, about 4% and 6% in the positive and negative loading phases, respectively. The maximum strength also occurs for around 40% lower top displacement demands. This can be due to some difficulty of the numerical model to simulate the beginning of the failure mechanism associated to the opening. For example, in the positive loading phase of the force–displacement curve of the experimental curve it is particularly notorious the strength degradation when compared to the response in the negative loading phase where the drop is softer. In the negative loading phase, the strength is underestimated by the numerical model, while, by contrast, it is slightly higher in the positive loading phase. Finally, the numerical ultimate strength is around 15% higher. About the capability of the numerical model to simulate the panel with a central door, Test DO2 (Fig. 7.13d), similar observations can be made, namely: – The initial stiffness is about 3% higher in the positive loading phase and 8% in the negative loading phase; – The numerical maximum strength is 2 and 18% higher in the positive and negative loading phases, respectively, and occurs for approximately the same top displacement; – The strength degradation is not totally well captured by the numerical model, since the ultimate strength is overestimated, namely 1.27 times and 1.41 times higher in the positive and negative loading phases, respectively.

7.4.3 OOP Behaviour The OOP behaviour has to be first calibrated independently and later with the interaction with the IP behaviour, to proceed with the definition of the IP-OOP displacement interaction. However, for that purpose, much more data is needed to calibrate that interaction adequately. As observed in Chap. 2, the IP-OOP interaction is not yet well characterized and some variability of the experimental results concerning the OOP strength for the same IP prior demand was found. Regarding the pure OOP behaviour, in Chap. 2 it was also observed that there is a need of more experimental data to assess the effect of different variables and to define possible range of results due to each one. Due to that, the calibration of the OOP behaviour and the calibration of the respective interaction with the IP demand was not performed. Therefore, in the next sub-sections, the simplified approach of considering a linear interaction between IP and OOP displacements and a linear elastic behaviour will be assumed. However, it is worth mentioning that, the assumed simplification revealed to be very efficient in a blind test prediction contest, where a 1:2.5 scaled three-storey RC infilled structure was first tested in a shaking-table and afterwards simulated. The test specimen is a 1:2.5 scale model composed of two parallel connected planar frames, with two bays and three storeys, making a structure with gross dimensions of 4.6 m length, 2.8 m width, and 3.9 m height. The RC structure was filled with masonry walls with openings in certain bays, as shown in Fig. 7.14.

292

7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

a)

b)

c)

Fig. 7.14 Blind test prediction contest: a lateral view of the specimen; b front view of the specimen and c general views

The specimen was subjected to ten unidirectional ground motions scaled from those recorded at the Herceg Novi station during the 15 April 1979 Montenegro earthquake with magnitude of 6.9. To account for the fact that the structure is constructed at √ 1:2.5 scale, the record was scaled in time by reducing the duration by a factor of 1/ 2.5. The record was base-line corrected and then scaled to match different levels of peak ground acceleration (0.05, 0.10, 0.20, 0.30, 0.40, 0.60, 0.70, 0.80, 1, and 1.2 g) that were used as input signals for the shaking table test. Nine international teams from different countries submitted their predictions, which included different modelling strategies such as detailed micro-modelling or simplified macro-modelling. The authors used different software applications and all the results were submitted to the organizational committee before the announcement of the winner. The final classification determined that the numerical model herein presented in this chapter was the most accurate one among all the submissions In fact, the numerical model captured well the collapse of the transverse panel from the ground-floor during the analysis, which was observed in the test [45]. This good demonstration of the numerical model potential allows to believe that this first and simplified proposal can simulate adequately the infills combined IP-OOP behaviour.

7.5 Case Study

293

As future works, the calibration of this combined IP-OOP behaviour has to be carried out based on the quasi-static tests performed and the shaking-table tests that were recently carried out in the Portuguese Laboratory of Civil Engineering.

7.5 Case Study The main aim of this sub-section is to present an application of the modelling approach presented in this Chapter. For this purpose, an eight-storey RC structure will be studied with three different configurations concerning the layout of the infill masonry walls. Despite the absence of calibration of the OOP and IP + OOP behaviour modelling, this preliminary study aims to assess the effect of the infill panels on the following topics: • Seismic vulnerability assessment of infilled RC structure (Parametric study 1); • Floor response spectra of infilled RC structure (Parametric study 2); • Seismic vulnerability assessment of damaged structure due to mainshockaftershock sequence (Parametric study 3). In order to accomplish this objective, several non-linear incremental dynamic analyses (IDA) were carried out to assess the seismic behaviour of the structure. The following sub-sections will present the description of the building structure under study as well all the considerations regarding the numerical modelling. Afterwards, each parametric study is presented independently in a specific sub-section, where each methodology is presented as well as the major results and findings.

7.5.1 Description of the Building The case study selected for this research work is an eight-storey building. The building has the plan dimensions of 20 m × 15 m, which consists of 4 × 5 m modules, with a storey height of 3 m (Fig. 7.15). The building was designed by the Portuguese Laboratory of Earthquake and Civil Engineering (LNEC) as part of a study on the seismic design of buildings, according to the design code enforced in the 80’s in Portugal [87]. The layout considered for the infill walls is illustrated in Fig. 7.15. The crosssection of the columns are: 30 × 60 cm (Storey 1 and 2); 30 × 50 cm (Storey 3 and 4); 40 × 30 cm (Storey 5 and 6) and 30 × 30 cm (Storey 7 and 8). The beams’ cross section is 30 × 60 cm with different reinforcement detailing according to the storey and layout in the structure. The reinforcement detailing of all the structural elements can be found in [87]. Concerning the design of the building, a global vertical load of 6.15kN/m2 plus a variable load of 2.5 kN/m2 was considered. A 3D model was generated in the software OpenSees [3], based on which three different numerical models were built considering different configurations

294

7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

Fig. 7.15 Case study: a Plant layout, b 3D bare frame model; c Front view; and d Lateral view

concerning the layout and strategies to simulate the infill walls, namely: (i) BF model: bare frame RC structure without infill masonry walls (only infills’ gravity load was considered); (ii) IP model: RC structure with infill walls distributed in the building’ façades considering only their non-linear in-plane behaviour; and (iii) IP_OOP model: RC structure with infill walls distributed similarly to the IP model, but considering their combined non-linear IP and OOP behaviour. The layout considered for the infill walls is illustrated in Fig. 7.15. Single leaf infill walls were selected for this study, representing the traditional unreinforced masonry panels constructed in Portugal, with full contact with the surrounding RC frame and composed of hollow clay weak bricks 150 mm thick. The relevant infill mechanical properties are described in Table 7.10. Here fc,h , fc,v are the values of infill panels compression strength parallel and normal to bed joints, while fw,u stands for the sliding shear strength of the mortar

7.5 Case Study

295

Table 7.10 Mechanical properties of the infill panels fc,h (MPa) fc,v (MPa) fw,u (MPa) Ss (MPa) Ei,h (MPa) Ei,v (MPa) G (MPa) W (kN/m3 ) 1.18

2.02

0.44

0.55

991

1873

1089

6.87

joints and Ss for the shear strength obtained from diagonal tensile strength test; Ew,h , Ew,v are the secant elastic modulus parallel and normal to bed joints, G is the shear modulus and W is the self-weight of the infill walls. The selected concrete grade is C25/30 and the steel reinforcement grade is A400 [87].

7.5.2 Numerical Modelling Strategy The numerical models were built in the OpenSees [3] based on the materials’ models and elements available in the software library. Different approaches can be assumed to simulate the seismic behaviour of RC frame elements, namely the use of elements with distributed inelasticity (force or displacement-based formulations) or elements with lumped-plasticity (with fixed length, i.e. the so-called plastic-hinge). For the present study, force-based beam-column elements were used to simulate the RC elements with fibre discretization that was adopted to simulate the behaviour at section level, where each fibre is associated with a uniaxial stress–strain law. The moment–curvature behaviour of the beam and column section is then obtained through the integration of the non-linear uniaxial stress–strain response of the individual fibres into which the section is subdivided. Regarding the infill walls simulation, it is used the equivalent double-strut model described in Sect. 8.3. The schematic layout of the modelling approach adopted to simulate the infilled RC frames is the same already shown in Fig. 7.1.

7.5.3 Material Modelling Properties This section presents the hysteretic models adopted for the concrete, reinforcing steel and infill walls and the values considered. For accurate modelling of the materials uniaxial stress–strain hysteretic rules, as required for consideration of the materials’ non-linearity, it is fundamental to find the best model that represents the real behaviour of the elements.

7.5.3.1

Concrete Modelling

The RC frame structure was modelled considering force-based BeamColumns elements in which the plastic hinge length (Lp ) of the RC elements were considered

296

7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

Table 7.11 Concrete mechanical parameters for the numerical model Compressive strength f c (MPa)

Compressive strain at peak strength εc (%)

Tensile strength f t (MPa)

25*

0.36

3.94

* Confinement

factor was considered and computed for each cross-section according to Mander et al. (1988) proposal

Table 7.12 Steel mechanical parameters for the numerical model Elastic modulus

Yield strength

Strain hardening parameter

Transition curve initial shape factor

Transition curve shape factors

Isotropic hardening factors

Es (GPa)

fy (MPa)

r (%)

R0

a1

a2

a3

a4

194.7

575.63

2.71

20

18.5

0.15

0.025

2

as half of their larger cross-section dimension. The modelling assumptions adopted for the beams and columns were based on the conclusions of the work developed by [46] where it was studied the biaxial flexural behaviour of RC columns. For the concrete modelling, the uniaxial material model Concrete02 was adopted, which is a uniaxial concrete material object with tensile strength and linear tension softening. More details concerning this uniaxial material model can be found in [47]. The adopted values are presented in Table 7.11.

7.5.3.2

Reinforcement Steel Modelling

For the steel reinforcement simulation, it was adopted the uniaxial model Steel 02 proposed by Menegotto et al. [88], which is commonly known as “uniaxial material Giuffre-Menegotto-Pinto” steel material object with isotropic strain hardening. The input properties for the steel reinforcement were a yield strength of 575.6 MPa, an elastic modulus equal to 194.7GPa, a strain-hardening ratio of 2.71%, a transition curve initial shape factor (Ro ) of 20 and the transition curve shape parameters a1 and a2 of 18.5 and 0.15 respectively. Finally, for the isotropic hardening parameters a3 and a4 the values 0.025 and 2 were assumed respectively. All the adopted values are summarized in Table 7.12.

7.5.3.3

Infill Masonry Walls’ Modelling

The uniaxial material model adopted to define the in-plane behaviour of the panel was adapted to the pinching 04 model, and the OOP behaviour was considered by putting the panel equivalent masses in the central nodes and with the element removal command. All the adopted values are summarized in Table 7.13. The results are in

7.5 Case Study

297

Table 7.13 Infill walls’ parameters for the numerical model fi,c (MPa)

dfi,c (%)

fi,y (MPa)

dfi,y (%)

fi,max (MPa)

dfi,max (%)

fi,u (MPa)

dfi,u (%)

0.57

0.1

0.75

0.25

1.01

0.5

0.28

1.5

IP drift (%) -1.5

-1.0

-0.5

0.0

0.5

1.0

1.5

2.0

0.10

3.3

0.08

2.7

0.06

2.0

0.04

1.3

0.02

0.7

0.00

0.0

-0.02

-0.7

-0.04

-1.3

-0.06

-2.0

-0.08

-2.7 -3.3

-0.10 -0.06

OOP drift (%)

OOP displacement (m)

-2.0

-0.04

-0.03

-0.01

0.00

0.01

0.03

0.04

0.06

IP displacement (m)

Fig. 7.16 In-plane and out-of-plane infill wall behaviour linear interaction adopted for the OOP model

agreement with the mechanical properties of infill panels made with hollow clay horizontal bricks. The IP and OOP drift limits adopted for the OOP model were set in order to follow a linear interaction between them, as illustrated in Fig. 7.16, with a maximum IP drift of 1.5% and OOP drift of 3%, the later corresponding to half-thickness of the wall. These values were selected based on previous research works [86].

7.5.4 Preliminary Results, Natural Frequencies and Vibration Modes From the eigenvalue analyses, the natural frequencies were extracted for each numerical model. In all models it was observed that the first vibration mode is pure translation in the longitudinal direction (herein also designated direction X), the second one is translation in the transverse direction (herein also designated direction Y) and finally the third one is a torsional mode. The three natural frequencies obtained by all the models are summarized in Table 7.14. Analytical prediction of the structure natural frequencies was carried out for the configuration without infill walls as described in [87].

298

7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

Table 7.14 Numerical results: Natural frequencies

2nd vibration mode (Hz)

3rd Vibration mode (Hz)

Analytical (BF 0.88 configuration)

1.06

1.39

0.89

1.07

1.46

Model

BF

1st vibration mode (Hz)

IP

1.94

2.03

5.73

IP-OOP

1.99

2.08

5.76

From the results it can be observed that the presence of the infill panels increased the natural frequencies about 2 times in the first two vibration modes and around 4 times in the third mode. The BF model first frequency is 0.89 Hz and for the IP and IP_OOP models it is 1.94 Hz and 1.99 Hz, respectively (2.18 and 2.24 times higher respectively). The BF model second natural frequency is 1.07 Hz compared with 2.03 Hz and 2.08 Hz achieved by the IP and IP_OOP models respectively. It can be also pointed that the BF model natural frequencies are quite close to those analytically predicted in [87].

7.5.5 Effect of the Infill Walls in the Seismic Vulnerability Assessment of Rc Buildings Parametric Study 1 7.5.5.1

Framework and Methodology

The need of retrofitting interventions or not for a specific building and the retrofitting objectives normally result from a detailed seismic assessment of the structure. However, the seismic assessment of existing buildings is a challenge due to numerous factors namely, knowledge level of the material properties, the clear identification of the structure type, etc. Eurocode 8 Part 3 [4] defines performance levels according to three different damage limit states: Damage Limitation (DL); Significant Damage (SD); and Near Collapse (NC). Typically, multiple performance criteria need to be satisfied. Therefore, prediction approaches based in vulnerability curves are being used by different authors for seismic vulnerability assessment of structures within a performance-based framework. Some studies can be found over the literature regarding the development of damage scales based in the concept of a single set of homogeneous vulnerability relationships. Different international codes and authors suggested damage state limits correlated to the maximum inter-storey drift ratio (ISDMAX ) achieved by the structure during the seismic analysis. For example, the HAZUS report [48] proposed four damage states for infilled RC structures (slight damage, moderate damage, extensive damage and collapse).

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On the other hand, VISION-2000 [49] proposed five limit states (fully operational, operational, life safe, near collapse, and collapse). Also, FEMA-273 [50] developed a damage scale based in six damage states, with the main difference relative to the VISION-2000 scale consisting in the drift limits for the immediate occupancy and damage control limit states. The remaining limit states are Life Safety, Limited Safety, Collapse Prevention and Collapse. Also EMS98 [51] proposed 5 different damage states: Grade 1, Grade 2, Grade 3, Grade 4 and Collapse, while the ATC-13 [52] scale is composed by six damages states: Slight, Light, Moderate, Heavy, Major and Collapse. Finally, Rosseto et al. [89] proposed a new damage scale (HCR-Scale) which was experimentally calibrated with ISDMAX , for structures with different lateral load resisting systems. The HCR-Scale is related to the existing damage scales described before and was adopted in the empirical curve generation with the objective of producing a single set of homogeneous vulnerability relationships to be applicable to all lateral-load resisting system. More details regarding the HCR-scale can be found in [89]. The HCR scale inter-storey drift limits and the corresponding expected typical damages are summarized in Table 7.15 and will be used in the present work. The seismic assessment was studied by plotting the fragility curves corresponding to each damage state. The material uncertainty was not considered for this study, since the main goal was to evaluate the RC structure’s seismic vulnerability, considering different modelling strategies to simulate the infill panels’ seismic behaviour. The damage limit state criteria based on the ISDMAX was adopted according to the Rosseto et al. [89] proposal. With the main purpose of deriving fragility curves for the infilled Table 7.15 HCR scale—Inter-storey drift limits for infilled RC frames according to [89] Damage state

Inter-storey drift (%) Typical damages expected

No damage

4.40

Complete or eminent building collapse

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7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour 1.0

3.0 2.5

Ground Accelaration (m2/)s

2.0 1.5 1.0

0.8

5

Probability of exceedance

Maximum inter-storey drift (%)

6

4

3

2

0.6

0.4

0.2

1

0.5

0.0 0.0

0.0 -0.5

0

-1.0

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-1.5 -2.0 -2.5 -3.0 0

10

20

30

40

50

60

70

80

90

0.2

0.3

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1.0

Pga (g)

Peak ground acceleration (g)

Time (sec)

Fragility curves

Incremental dynamic analyses 20 records Pga=[0.05-1g]

ISDMAX

Slight, Light, Moderate, Extensive, Partial Collapse and Collapse

Fig. 7.17 Seismic vulnerability assessment: methodology

structures, a framework was developed to handle the various inputs and outputs of the non-linear analyses performed (Fig. 7.17). The overall process can be summarized in the following steps: (1) execution of non-linear incremental dynamic analyses (IDA) for each numerical model using a set of progressively scaled ground motion records, with the main objective of extracting the ISDMAX results; (2) association of each numerical model damage state based on the criteria presented in Table 7.15; (3) calculation of the cumulative percentage of outputs that reached each damage limit state demand according to the pga; (4) derivation and plotting fragility curves for the corresponding damage states (S-slight, L-light, M-moderate, E-extensive, PC-partial collapse and C-collapse) for each numerical model. The earthquake records for these analyses were generated using a set of 20 ground motions that were selected from real previous seismic events according to the [53] method. The ground-motions were divided into 10 Type 1 and 10 Type 2 records, according to Eurocode 8 [4], for seismic zone 1.3 (seismic action type 1) and 2.3 (seismic action type 2) and ground type A. Bidirectional ground-motions were applied in each analysis, for which the corresponding elastic spectra are plotted in Fig. 7.18.

7.5.5.2

Numerical Results and Discussion

More than 200 analyses were carried out to evaluate the performance of the three numerical models (BF, IP and IP_OOP) and the influence of considering the OOP effect of the infill panels. Figure 7.19 and Fig. 7.20 show plots of the evolutions of the maximum inter-storey drift (ISDMAX ) where, the drift thresholds for each damage state are also plotted as horizontal dashed lines identified by the right scale. It can be observed that the BF model has higher maximum inter-storey drifts for peak ground acceleration above 0.2 g and that the transverse direction (direction Y) is the most vulnerable direction of the building, for which the maximum inter-storey drift increases quickly after 0.14 g.

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301

Type 1

EC8 Elastic Spectra ground-motion 1 ground-motion 2 ground-motion 3 ground-motion 4 ground-motion 5 ground-motion 6 ground-motion 7 ground-motion 8 ground-motion 9 ground-motion 10 Average

0.8

Sd (g)

0.6

0.4

0.2

0.0 0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

Period (s)

a) 1.0

Type 2

EC8 Elastic Spectra ground-motion 11 ground-motion 12 ground-motion 13 ground-motion 14 ground-motion 15 ground-motion 16 ground-motion 17 ground-motion 18 ground-motion 19 ground-motion 20 Average

0.8

Sd (g)

0.6

0.4

0.2

0.0 0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

Period (s)

b) Fig. 7.18 Elastic spectra of the ground motions used for the IDA analyses: a Type 1; b type 2

A significant difference is observed between the IP model and the IP_OOP model: (i) for the IP model the infills are protective for peak ground acceleration until 0.2 g and with similar building response in both directions for peak ground acceleration beyond 0.2 g; (ii) the IP_OOP model shows a similar response to that of the BF model, but with higher maximum inter-storey drifts for peak ground acceleration of 0.17 g in the longitudinal direction (Direction X) and 0.1 g in the transverse direction (Direction Y). The evolutions of the maximum base shear are plotted in Fig. 7.21 and it can be observed that the IP and IP_OOP models (Fig. 7.21b, c respectively) reached about two times higher base shear values than the BF model, which is due to the presence

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7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

Fig. 7.19 Incremental dynamic analysis results: ISDMAX (Direction X) for a BF model; b IP model; and c IP_OOP model

of the infill walls. This significant increase of the maximum base shear can cause major damage and even building collapse if the foundations are not designed for such seismic loading. Concerning the transverse direction, the conclusions were similar. Figure 22a shows the results of one record that was scaled at three levels for reaching the peak ground acceleration of: (i) 0.15 g; (ii) 0.3 g; and (iii) 0.5 g, as an example. For this ground motion it was found that for the peak ground acceleration of 0.15 g, the BF model reached the highest ISD (Fig. 7.22a). The BF model shows four times higher drift for the peak ground acceleration of 0.30 g (Fig. 7.22b), namely for drifts between storeys 0–1 and 2–3. The IP and IP_OOP models exhibit slight differences and the latter shows two times higher drift values between storeys 0–1 and 2–3. It can be observed that for the peak ground acceleration of 0.5 g (Fig. 7.22c) the most vulnerable model is the IP_OOP, with the collapse of storeys 1 and 5. Is also noted that in the BF model the 1st and 2nd storeys reach collapse and, finally, in the IP model no significant drift values are observed for both directions, the most vulnerable storey being the 3rd one with a maximum drift of 1.5%. It is clearly important to consider the OOP

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Fig. 7.20 Incremental dynamic analysis results: ISDMAX (Direction Y) for a BF model; b IP model; and c IP_OOP model

behaviour of the infill walls when subjected to earthquakes, since the increasing IP drift values cause a decrease of the OOP capacity of the infill wall and finally result in the collapse of these non-structural elements. It was observed that the infill walls’ collapse may introduce a vertical irregularity and soft-storey mechanism may develop in the buildings. This response, as observed in the last decade in real earthquakes, is likely to result in the buildings’ collapse. From the obtained fragility curves (see Fig. 7.23), it can be observed that for the slight and light damage levels all three numerical models reached similar peak ground acceleration levels of 0.05–0.12 g. The moderate damage state is reached first in the BF model (Fig. 7.23a) at about 0.15 g and the IP and IP_OOP models at 0.25– 0.3 g (Fig. 7.23b, c). Finally, it is observed that for large peak ground acceleration demands, the IP_OOP model reached the moderate and collapse damage states for lower values than the other numerical models. To better understand the difference between the three numerical models’ performances, the moderate, extensive and collapse damage states were separated and plotted as shown in Fig. 7.24a, b and c, respectively. Concerning moderate damage, it is observed that the BF model is the most vulnerable one and that the infill walls with only IP behaviour positively contribute to protect the building.

304

7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour 7000

7000

20 IDA curves Longitudinal Direction

6000

Maximum base shear (kN)

Maximum base shear (kN)

6000 5000 4000 3000 2000

4000 3000 2000 1000

1000 0 0.0

5000

0.1

0.2

0.3

0.4

0.5

0.6

0.7

Longitudinal Direction IP model

0 0.0

0.8

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

pga (g)

pga (g)

a)

b) 7000

Maximum base shear (kN)

6000 5000 4000 3000 2000 1000 0 0.0

Longitudinal Direction IP_OOP model 0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

pga (g)

c) Fig. 7.21 Incremental dynamic analysis results: maximum base shear evolution of a BF, b IP, and c IP_OOP models, for the longitudinal direction (20 IDA curves)

Fig. 7.22 Incremental dynamic analysis results: envelope of ISDMAX : a pga = 0.15 g, b pga = 0.30 g, and c pga = 0.50 g

305

1.0

1.0

0.8

0.8

Probability of exceedance

Probability of exceedance

7.5 Case Study

0.6

0.4 Slight Light Moderate Extensive Part. Collapse Collapse

0.2

0.0 0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

0.6

0.4

0.0 0.0

1.0

Slight Light Moderate Extensive Part. Collapse Collapse

0.2

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1.0

Pga (g)

Pga (g)

a)

b) 1.0

Probability of exceedance

0.8

0.6

0.4 Slight Light Moderate Extensive Part. Collapse Collapse

0.2

0.0 0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1.0

Pga (g)

c)

Fig. 7.23 Fragility curves for a BF model, b IP model, and c IP_OOP model

Extensive damage occurs earlier in the BF model for lower peak ground acceleration values (0.4 g. In terms of the collapse damage state, it is observed that the OOP behaviour of the infill walls is critical. In fact, clear differences are observed between the IP_OOP model and the BF and IP models, which reinforces the need to consider the OOP behaviour of the infill walls in the seismic safety assessment of the existing buildings, consequently, in the numerical models.

7.5.5.3

Summary of the Major Conclusions of Parametric Study 1

The parametric study 1 was carried out aiming to assess the effect of the infill panels in the seismic vulnerability assessment of RC structures. For this, non-linear dynamic analyses were performed to collect the maximum inter-storey drift ratios (ISDMAX ) and compared with drift limits defined for specific damage states, namely: slight, light, moderate, partial collapse and collapse. Three different assumptions were assumed concerning the modelling of the infill panels, namely: (i) bare frame configuration (without infill panels), (ii) infilled RC frame configuration considering only the infills IP behaviour, and (iii) infilled RC frame configuration considering the infills IP-OOP behaviour.

7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour 1.0

1.0

0.8

0.8

Probability of exceedance

Probability of exceedance

306

0.6

0.4

BF IP IP_OOP Damage state: moderate

0.2

0.0 0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

0.6

0.4

0.0 0.0

1.0

BF IP IP_OOP Damage state: Extensive

0.2

0.1

0.2

0.3

0.4

Pga (g)

0.5

0.6

0.7

0.8

0.9

1.0

Pga (g)

a)

b)

Probability of exceedance

1.0

BF IP IP_OOP 0.8 Damage state: Collapse 0.6

0.4

0.2

0.0 0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1.0

Pga (g)

c)

Fig. 7.24 Fragility curves for a moderate, b extensive and c collapse damage states

From the study, it was first observed that the presence of the infill panels increased the natural frequencies of the RC structure about 2 times. It was observed that the presence of the infills modify the expected response of the building when subjected to an earthquake. Concerning the base shear, the infills contributed to the increase of the base shear around 2 times, which should be an issue correctly analysed by a structural designer during the assessment of an infilled RC structure (or design of a new one). The incorrect evaluation of the expected base shear could result in serious consequences to the building structure in an earthquake scenario. It was observed that, for lower seismic demands, the models with infill panels reached lower ISDMAX when compared with the BF model response. For medium– high seismic demands, the infills IP-OOP behaviour played an important role, since it contributed to the substantial increase of the ISDMAX due to the collapse of panels. This phenomenon occurred because many panels reached the IP-OOP drift limits and collapsed during the earthquake, triggering the development of soft-storey mechanism in some cases. The ISDMAX reached by the model considering the IPOOP behaviour were higher than those reached by the bare frame configuration were. Concerning the model in which was considered only the infills IP behaviour, it was observed that for low seismic demands the level of displacements, however with the

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307

increase of the seismic demand there is a sudden increase of the ISDMAX which is due to the infills deterioration. This fact highlights the major conclusion of this parametric the study, which is the fact that the combined IP-OOP behaviour of the infill walls should be considered during a seismic assessment of an infilled RC structure. At least, the IP stiffness and strength contribution to the building structure response should be considered. The past and actual procedures, disregarding the contribution of the infill panels in the seismic response of RC buildings response, should be revised in future updates of Eurocode 8.

7.5.6 Effect of the Infill Panels in the Floor Response Spectra Parametric Study 2 7.5.6.1

Objectives and Methodology

The present section will study the effect of the infill panels in the floor seismic response spectra of RC buildings, for which it will be considered only the BF and IP_OOP models. The first objective of this parametric study is to assess the infill panels’ influence in the global seismic response of the structure, through the analysis of the average values of floor accelerations and velocities extracted from all the analyses carried out, which allowed determining the peak floor acceleration (PFA) and the peak floor velocity (PFV). At the same time, floor response spectra (FRS) were drawn for the design pga according to Eurocode 8 [4] which, in the present case, corresponds to 1.5 m/s2 . The FRS of the infill model is compared with the BF model FRS, which will allow to assess any modification of the spectra shape and peaks due to the infills presence. Capacity response spectra (CRP) are also obtained and presented according to the Eurocode 8 [4] proposal. The second goal of this parametric study is the study of the seismic response of the infill panels themselves. For this purpose, the OOP accelerations and velocities measured in the centre of the panel were recorded for all the input ground-motions. Maximum OOP acceleration and velocity values reached by the infill panels at each floor were obtained and derived, respectively by PIA and PIV. The respective infill acceleration and capacity response spectra (IARS and IACS) and infill velocity response spectra (IVRS) were then derived and are also presented. Finally, the in-elevation amplification of those values will be evaluated and compared with acceleration demands suggested by Eurocode 8. The schematic workflow layout of this methodology is presented in Fig. 7.25. Incremental dynamic analyses were carried out to extract those results. For the IDA, it was used the set of 10 bi-directional ground motions, type 1, presented in Sect. 7.5.3.1. Again, it is important to remember that this parametric study will only

308

7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

Incremental Dynamic Analyses

Infill panels analysis

Global Analysis

ISDMAX

PFA

FRS

CRS

PFV

PIA

IARS

Study of the infill masonry walls influence ISDMAX –Maximum inter-storey drift ratio PFA –Peak floor acceleration FRS –Floor response spectra CRS –Capacity response spectra PFV –Peak floor velocity

EC8

IACS

PIV

IVRS

PIA –peak infill acceleration IARS –infill acceleration response spectra IACS –Infill acceleration capacity response spectra PIV –Peak infill velocity IVRS –Infill velocity response spectra

Study of the infill masonry walls seismic behavior

Fig. 7.25 Schematic layout of the workflow

study the results of the BF and IP_OOP models, since the effect of considering the infills combined IP-OOP behaviour was studied in Parametric study 1.

7.5.6.2

Analysis of the Global Structural Seismic Behaviour

The PFA results are shown in Fig. 7.26 from which the following observations can be drawn: (i) larger variation is found in the longitudinal direction (from 2 to 2.8 times higher) for the BF model and in the transverse direction for the IP_OOP model (from 1.87 times to 3.2 times higher), respectively in Fig. 7.26a, d; (ii) variation of the PFA between 0.05 g and 1.45 g and between 0.03 g and 2.8 g, respectively in the BF and IP_OOP model; (c) from the comparison between the BF and IP_OOP average curves (Fig. 7.26e, f) it is possible to observe that the infills globally may induce an increment of the PFA, more pronounced along the transverse direction where the it reached from 22 to 66% higher PFA when compared to the 25%-48% reached in the longitudinal direction. The PFA/pga ratio plots are shown in Fig. 7.27 and allow to understand the evolution of the PFA with the increment of pga demand. In both models it appears that the PFA/pga ratio tends to decrease with the increase of the pga and tends to 1 for larger pga values in the case of the BF model (Fig. 7.27a, b). The main difference is that the IP_OOP model tends to decrease to values around 2 with the increase of the pga, and seems to have the same decrease trend in both the longitudinal and transverse direction (Fig. 7.27c, d). From the comparison between the BF and IP_OOP model, it is possible to observe similar responses in both building directions, namely the decrease of the PFA/pga ratio from 3–3.5 to 2 and 2.5 to 1.3 in the case of the

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3.0

3.0

BF model Transverse direction

2.5

2.5

2.0

2.0

PFA (g)

PFA (g)

BF model Longitudinal direction

1.5

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pga (g)

a)

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PFA (g)

0.4

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IP_OOP model Longitudinal direction

2.5

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pga (g)

d)

Longitudinal direction BF model IP_OOP model

1.4 1.2 1.0

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0.8

PFA (g)

1.0

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pga (g)

e)

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c)

PFA (g)

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IP_OOP model Longitudinal direction

1.2

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b)

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1.4

0.3

pga (g)

0.0 0.0

0.1

0.2

0.3

0.4

0.5

pga (g)

f)

Fig. 7.26 Global results of PFA: a BF model (Longitudinal); b BF model (transverse); c IP_OOP model (Longitudinal); d IP_OOP model (transverse); e Comparison (longitudinal) and f Comparison (transverse)

IP_OOP and BF model, respectively (see Fig. 7.27e, f). Globally, it is impossible to identify the infill panels’ contribution regarding the relationship between the PFA and the seismic demand, since there are similar trends. The main difference is related to the PFA/pga ratio, which is higher in the case of the IP_OOP model. The storey PFA heightwise profiles are plotted in Fig. 7.28 for four levels of pga, namely 0.08, 0.25, 0.37 and 0.50 g. From these plots it is possible to observe that the infill panels tend to increase the PFA for larger pga demands and that the maximum

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7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

10

10

BF model Longitudinal direction

9

8

8

7

7

Ratio PFA/pga

Ratio PFA/pga

9

6 5 4 3

6 5 4 3

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BF model Transverse direction

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a)

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7

7

Ratio PFA/pga

Ratio PFA/pga

9

6 5 4 3

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IP_OOP model Transverse direction

6 5 4 3 2

2

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c)

d) 4.0

Longitudinal direction

BF model IP_OOP model

3.5

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Ratio PFA/pga

3.0

2.5 2.0 1.5

1.5 1.0

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0.2

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pga (g)

e)

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BF model IP_OOP model

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Transverse direction

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pga (g)

pga (g)

Ratio PFA/pga

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b)

IP_OOP model Longitudinal direction

8

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10

10 9

0.3

pga (g)

pga (g)

0.0 0.0

0.1

0.2

0.3

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0.5

pga (g)

f)

Fig. 7.27 Global results: Ratio PFA/pga a BF model (Longitudinal); b BF model (transverse); c IP_OOP model (Longitudinal); d IP_OOP model (transverse); e Comparison (longitudinal) and f Comparison (transverse)

PFA values were reached at the third and fourth storey for the longitudinal direction and at the first, second and third storey for the transverse direction. It is not observed the matching with the storeys where the maximum PFA values occurred in the BF model, which supports the idea that the infill panels can change the location of the floor where the maximum PFA values occur. Finally, it can be mentioned that larger amplification of the PFA due to the infills is observed in the transverse direction and that for a pga equal to 0.08 g the PFA profiles are similar. Curiously, it is observed a

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8

7

6

6

5

BF model

Storey level

Storey level

8

Longitudinal direction

7

PFA BF,average,0.08g

4

PFA BF,average,0.25g PFA BF,average,0.37g PFA BF,average,0.50g

3

Transverse direction

5

BF model

4

PFA BF,average,0.25g

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PFA BF,average,0.08g PFA BF,average,0.37g

IP_OOP model 2

IP_OOP model

PFA IP_OOP,average,0.08g

2

PFA IP_OOP,average,0.08g

1

PFA IP_OOP,average,0.37g

PFA IP_OOP,average,0.25g

PFA IP_OOP,average,0.25g

1

PFA IP_OOP,average,0.37g PFA IP_OOP,average,0.50g

0 0.0

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0 0.0

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PFA (g)

PFA (g)

a)

b)

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1.4

Fig. 7.28 Global results: PFA heightwise profiles a Longitudinal; b Transverse

continuous, yet but slight, in-elevation increase of the PFA in the BF model, which is not observed in the IP_OOP model results. The floor acceleration response spectra (FRS), for 5% damping, for each storey and each model, in the longitudinal direction, are plotted in Fig. 7.29, from which it is possible to observe different shapes and different peaks in each floor spectra. From the comparison of spectra it can be seen in some storeys the existence of two peaks for periods between 0.05 and 0.40 s in the IP_OOP model response, instead of two peaks for periods in the range 0.4–1.1 s for the BF response. In the top storeys (5 to 8), the shape of the curves are very similar despite the higher pga demand reached by the IP_OOP model, and as mentioned before the shift of the peaks for lower periods. Finally, it is also noticed that the FRS change throughout the building height, which can be related to the higher modes of the structure. Capacity response spectra (CRS) were plotted based in the Eurocode 8 [4] proposal, again for each storey, each numerical model and only referring to the longitudinal direction (Fig. 7.30). Obviously, similar trends can be found in these results, namely the difference in terms of shape and modification of the CRS peaks and also the in-elevation modification of the CRS curves. From the results, it can be noticed that the BF model reaches higher displacements demands (Sd ) than the IP_OOP model (for storey 1, 2 and 3), respectively 3, 2 and 1.2 times higher. Concerning the remaining floors, the peak displacement demands are similar in both models, but the main difference is related to the higher acceleration demands (Sa ) reached by the IP_OOP model. These results show, again, the influence of the infill walls in the CRS of the building structure. The building peak floor velocity during a seismic event is a topic not covered over the literature, which may be revealed with the fact that the floor accelerations are more directly related to the seismic forces and damages that are developed within the structural elements. Nonetheless, in the present study the floor velocities were recorded for all the analyses (Fig. 7.31) and some observations can be drawn: (i) the average PFV of the

312

7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

3.0

3.0

ACC Average BF model Longitudinal Direction Storey level: 1

2.5

2.0

Sa (g)

Sa (g)

Sa (g)

1.5

1.5

1.0

1.0

0.5

0.5

0.5

0.0 0.0

0.0 0.0

1.0

1.5

2.0

2.5

3.0

3.5

4.0

0.5

1.0

1.5

3.0

3.5

0.0 0.0

4.0

ACC Average IP_OOP model Longitudinal Direction Storey level: 2

1.0

0.5

0.5

0.5

0.0 0.0

0.0 0.0

2.0

2.5

3.0

3.5

4.0

0.5

1.0

1.5

2.5

3.0

3.5

0.0 0.0

4.0

0.5

0.5

1.0

1.5

2.0

2.5

3.0

3.5

0.0 0.0

4.0

1.0

1.5

2.0

2.5

3.0

3.5

ACC Average IP_OOP model Longitudinal Direction Storey level: 4

Sa (g)

Sa (g)

1.0

0.5

0.0 0.0

0.0 0.0

3.0

3.5

4.0

0.5

1.0

1.5

Period(s)

2.0

2.5

3.0

3.5

0.0 0.0

4.0

Sa (g)

Sa (g)

1.0

0.5

0.5

0.5

0.0 0.0

0.0 0.0

3.0

3.5

4.0

0.5

1.0

1.5

Period(s)

2.0

2.5

3.0

3.5

0.0 0.0

4.0

Sa (g)

Sa (g)

1.0

0.5

0.5

0.5

0.0 0.0

0.0 0.0

2.5

3.0

3.5

4.0

3.0

3.5

4.0

BF model IP_OOP model Longitudinal Direction Storey level: 5

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

BF model IP_OOP model Longitudinal Direction Storey level: 6

1.5

1.0

2.0

2.5

2.0

1.5

1.0

Period(s)

2.0

2.5

2.0

1.5

1.5

1.5

3.0

ACC Average IP_OOP model Longitudinal Direction Storey level: 6

2.5

2.0

1.0

1.0

Period(s)

3.0

ACC Average BF model Longitudinal Direction Storey level: 6

2.5

0.5

0.5

Period(s)

3.0

4.0

1.5

1.0

2.5

3.5

2.0

1.5

1.0

2.0

3.0

BF model IP_OOP model Longitudinal Direction Storey level: 4

2.5

2.0

1.5

2.5

3.0

ACC Average IP_OOP model Longitudinal Direction Storey level: 5

2.5

2.0

2.0

Period(s)

3.0

ACC Average BF model Longitudinal Direction Storey level: 5

2.5

1.5

1.5

Period(s)

3.0

1.0

1.0

1.5

0.5

2.5

4.0

2.0

1.5

0.5

0.5

0.5

2.5

1.0

2.0

3.5

BF model IP_OOP model Longitudinal Direction Storey level: 3

3.0

1.0

1.5

3.0

1.5

0.0 0.0

4.0

2.0

1.5

2.5

Period(s)

2.5

2.0

2.0

0.5

0.5

3.0

ACC Average BF model Longitudinal Direction Storey level: 4

2.5

1.0

1.5

Period(s)

3.0

4.0

1.0

Period(s)

0.5

1.0

2.5

Sa (g)

Sa (g)

1.0

3.5

2.0

1.5

1.0

0.5

0.5

3.0

ACC Average IP_OOP model Longitudinal Direction Storey level: 3

2.0

1.5

3.0

Period(s)

2.5

2.0

0.0 0.0

2.0

3.0

ACC Average BF model Longitudinal Direction Storey level: 3

2.5

2.5

BF model IP_OOP model Longitudinal Direction Storey level: 2

Period(s)

Period(s) 3.0

2.0

1.5

1.0

1.5

1.5

2.0

1.5

1.0

1.0

1.0

2.5

Sa (g)

1.5

0.5

0.5

3.0

2.0

Sa (g)

Sa (g)

3.0

Period(s)

2.5

2.0

Sa (g)

2.5

3.0

ACC Average BF model Longitudinal Direction Storey level: 2

2.5

Sa (g)

2.0

Period(s)

Period(s)

Sa (g)

1.5

1.0

0.5

BF model IP_OOP model Longitudinal Direction Storey level: 1

2.5

2.0

2.0

Sa (g)

3.0

ACC Average IP_OOP model Longitudinal Direction Storey level: 1

2.5

0.5

1.0

1.5

2.0

Period(s)

2.5

3.0

3.5

4.0

0.0 0.0

0.5

1.0

1.5

2.0

Period(s)

Fig. 7.29 Global results: FRS a BF model; b IP_OOP model; and c Comparison

2.5

3.0

3.5

4.0

7.5 Case Study

313

3.0

3.0

ACC Average BF model Longitudinal Direction Storey level: 7

2.5

2.0

Sa (g)

Sa (g)

Sa (g)

1.5

1.5

1.5

1.0

1.0

1.0

0.5

0.5

0.5

0.0 0.0

0.0 0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

0.0 0.0

4.0

3.0

0.5

0.5

2.0

2.5

Period(s)

a)

3.0

3.5

4.0

Sa (g)

Sa (g)

1.0

0.0 0.0

2.5

3.0

3.5

4.0

BF model IP_OOP model Longitudinal Direction Storey level: 8

2.0

1.5

1.0

2.0

2.5

2.0

1.5

1.5

1.5

3.0

ACC Average IP_OOP model Longitudinal Direction Storey level: 8

2.5

2.0

1.0

1.0

Period(s)

3.0

ACC Average BF model Longitudinal Direction Storey level: 8

2.5

0.5

0.5

Period(s)

Period(s)

0.0 0.0

BF model IP_OOP model Longitudinal Direction Storey level: 7

2.5

2.0

2.0

Sa (g)

3.0

ACC Average IP_OOP model Longitudinal Direction Storey level: 7

2.5

1.5

1.0

0.5

0.5

1.0

1.5

2.0

Period(s)

b)

2.5

3.0

3.5

4.0

0.0 0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

Period(s)

c)

Fig. 7.29 (continued)

BF model ranges from 0.21 to 0.59 m/s and 0.24 to 0.49 m/s, respectively, in the longitudinal and transverse directions, which is curious since it was observed higher PFA in the transverse direction (Fig. 7.31a, b); (ii) concerning the infill model, the average PFV curve ranges from 0.20 to 0.69 m/s and 0.23 to 0.62 m/s, respectively in the longitudinal and transverse direction; again, it is noticed that larger values were reached in the longitudinal response (Fig. 7.31c, d); (iii) the comparison between the BF model and the infill model (Fig. 7.31e, f) shows that along the longitudinal direction both models have similar trend with slight differences for pga equal to 0.3 g where the infill model reaches 12% higher PFV. Concerning the transverse direction, the PFV reached by the BF model are higher than those obtained by the infill model (from 1.1 to 1.6 times higher) for pga values lower than 0.25 g. For pga values greater than 0.25 g, the infill model reaches values 20% higher than those obtained in the BF model. Finally, it can be mentioned that both model present an increase of the PFV until pga equal to 0.3 g and then tend to be more stable for pga values until 0.5 g. To finish the assessment of the effect of the infill panels’ presence in the global dynamic behaviour of the structure, the PFV heightwise profiles are presented in Fig. 7.32, again for four different seismic demands (0.08, 0.25, 0.37 and 0.50 g). From the analysis of those results, it can be noticed that similar trends can be observed in both directions of the two models. Besides the continuous in-elevation increase of the PFV, different responses can be observed concerning the BF and the IP_OOP model, namely: (a) higher PFV values reached by the BF model in all the storeys, for pga equal to 0.08 g, along the longitudinal direction (see Fig. 7.32a); the differences between the BF and IP_OOP models responses concerning the PFV is approximately constant in each storey; (b) for a seismic demand equal to 0.25 and 0.37 g, the BF model reached higher PFV than that of IP_OOP model for the storey 1, 2 and 3

314 2.1

ACC Average BF model Longitudinal Direction Storey level: 1

2.1 1.8

ACC Average IP_OOP model Longitudinal Direction Storey level: 1

2.1 1.8

1.5

1.5

1.2

1.2

1.2

0.9

0.9

0.6

0.6

0.3

0.3

0.0 0.0

0.1

0.2

0.3

0.4

0.5

0.0 0.0

0.6

Sa(g)

1.5

Sa(g)

Sa(g)

1.8

7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

0.3

0.1

0.2

ACC Average BF model Longitudinal Direction Storey level: 2

2.1 1.8

2.1 1.8

1.2

0.9

0.9

0.6

0.6

0.3

0.3

0.2

0.3

0.4

0.5

0.0 0.0

0.6

Sa(g)

1.5

1.2

0.1

2.1 1.8

0.2

0.3

0.4

0.5

0.0 0.0

0.6

2.1 1.8

0.9

0.3

0.2

0.3

0.4

0.5

0.0 0.0

0.6

Sa(g)

1.2

Sa(g)

1.5

0.3

1.8

2.1 1.8

0.2

0.3

0.4

0.5

0.0 0.0

0.6

ACC Average IP_OOP model Longitudinal Direction Storey level: 4

2.1 1.8

1.2

0.9 0.6

0.3

0.3

0.2

0.3

0.4

0.5

0.0 0.0

0.6

Sa(g)

1.5

1.2

Sa(g)

1.5

0.6

1.8

2.1 1.8

0.2

0.3

0.4

0.5

0.0 0.0

0.6

ACC Average IP_OOP model Longitudinal Direction Storey level: 5

2.1 1.8

0.9 0.6

0.3

0.3

0.2

0.3

0.4

0.5

0.0 0.0

0.6

Sa(g)

1.2

Sa(g)

1.5

1.2

0.6

1.8

2.1 1.8

0.2

0.3

0.4

0.5

0.0 0.0

0.6

ACC Average IP_OOP model Longitudinal Direction Storey level: 6

2.1 1.8

0.9 0.6

0.3

0.3

0.2

0.3

Sd(m)

0.4

0.5

0.6

Sa(g)

1.2

Sa(g)

1.5

1.2

0.1

0.2

0.0 0.0

0.3

0.4

0.5

0.6

0.4

0.5

0.6

0.4

0.5

0.6

BF model IP_OOP model Longitudinal Direction Storey level: 5

0.1

0.2

0.3

Sd(g)

1.5

0.0 0.0

0.1

0.3

0.1

1.2

0.6

0.6

0.9

1.5

0.9

0.5

BF model IP_OOP model Longitudinal Direction Storey level: 4

Sd(m)

ACC Average BF model Longitudinal Direction Storey level: 6

0.4

0.6

Sd(m) 2.1

0.3

Sd(g)

1.5

0.1

0.2

0.3

0.1

1.2 0.9

0.6

0.9

1.5

0.0 0.0

0.1

Sd(m)

ACC Average BF model Longitudinal Direction Storey level: 5

0.5

0.6

Sd(m) 2.1

0.4

Sd(g)

1.2 0.9

0.3

0.3

0.1

Sd(m)

ACC Average BF model Longitudinal Direction Storey level: 4

0.1

0.2

0.9

1.5

0.0 0.0

0.1

BF model IP_OOP model Longitudinal Direction Storey level: 3

0.6

Sd(m) 2.1

0.6

Sd(g)

ACC Average IP_OOP model Longitudinal Direction Storey level: 3

1.2

0.6

0.5

0.3

0.1

1.2

0.6

0.4

0.6

1.5

0.9

0.3

BF model IP_OOP model Longitudinal Direction Storey level: 2

Sd(m)

ACC Average BF model Longitudinal Direction Storey level: 3

0.1

0.2

0.9

1.5

0.0 0.0

0.1

Sd(g)

1.5

1.8

Sa(g)

0.0 0.0

0.6

1.2

2.1

Sa(g)

0.5

ACC Average IP_OOP model Longitudinal Direction Storey level: 2

Sd(m)

Sa(g)

0.4

1.5

0.0 0.0

Sa(g)

0.3

Sd(m)

Sa(g)

Sa(g)

1.8

0.9 0.6

Sd(m) 2.1

BF model IP_OOP model Longitudinal Direction Storey level: 1

BF model IP_OOP model Longitudinal Direction Storey level: 6

0.9 0.6 0.3

0.1

0.2

0.3

0.4

0.5

0.6

0.0 0.0

0.1

0.2

Sd(m)

Fig. 7.30 Global results: CRS a BF model; b IP_OOP model; and c Comparison

0.3

Sd(g)

7.5 Case Study 2.1

ACC Average BF model Longitudinal Direction Storey level: 7

2.1 1.8

ACC Average IP_OOP model Longitudinal Direction Storey level: 7

2.1 1.8

1.5

1.5

1.2

1.2

1.2

0.9

0.9

0.6

0.6

0.3

0.3

0.0 0.0

0.1

0.2

0.3

0.4

0.5

0.0 0.0

0.6

Sa(g)

1.5

Sa(g)

Sa(g)

1.8

315

0.3

0.1

0.2

0.3

0.4

0.5

0.0 0.0

0.6

ACC Average BF model Longitudinal Direction Storey level: 8

2.1 1.8

ACC Average IP_OOP model Longitudinal Direction Storey level: 8

2.1 1.8 1.5

1.2

1.2

0.9

0.9

0.6

0.6

0.3

0.3

0.2

0.3

0.4

Sd(m)

a)

0.5

0.6

Sa(g)

1.5

1.2

0.1

0.2

0.0 0.0

0.3

0.4

0.5

0.6

0.4

0.5

0.6

Sd(g)

1.5

0.0 0.0

0.1

Sd(m)

Sa(g)

Sa(g)

1.8

0.9 0.6

Sd(m) 2.1

BF model IP_OOP model Longitudinal Direction Storey level: 7

BF model IP_OOP model Longitudinal Direction Storey level: 8

0.9 0.6 0.3

0.1

0.2

0.3

0.4

0.5

0.6

Sd(m)

b)

0.0 0.0

0.1

0.2

0.3

Sd(g)

c)

Fig. 7.30 (continued)

(longitudinal direction); concerning the upper storeys (4 to 8), the highest PFV was reached by the IP_OOP model; c) for the pga 0.50 g, the IP_OOP model reached the highest PFV in all the building storeys (longitudinal direction); d) concerning the transverse direction, for the lowers seismic demands (0.08 and 0.25 g), again the BF model reached the highest PFV, however the differences between the BF model and IP_OOP model PFV reduces from the bottom to the upper storeys; e) the highest PFV in the storeys 1 and 2, in the longitudinal direction, is always reached by the IP_OOP model. On the other hand, in the remaining storeys (3, 4, 5, 6, 7 and 8) the highest PFV is reached by the IP_model, as shown in Fig. 7.32b. To conclude, it seems that for lower seismic demands, the model without infill panels (BF) achieves the highest PFV, and for medium–high seismic demands the highest PFV is achieved by the IP_OOP model.

7.5.6.3

Analysis of the Infill Walls Seismic Behaviour

This sub-section presents the study of the seismic behaviour of the infill panels during the seismic event, which is of particular interest concerning the development of assessment methodologies of the seismic vulnerability of those elements when subjected to OOP accelerations. Thus, as previously mentioned this assessment focuses on the OOP accelerations and velocities that the panels are subjected to, during and earthquake, taking advantage of the numerical strategy adopted to model the seismic behaviour of the infill walls, that allows simulating the combined in-plane and out-of-plane behaviour. It is important to remember that this numerical model allows recording the displacements, velocities and accelerations of the central nodes.

316

7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour 1.50

1.50

BF model Transverse direction

1.25

1.25

1.00

1.00

PFV (m/s)

PFV (m/s)

BF model Longitudinal direction

0.75

0.75

0.50

0.50

0.25

0.25

0.00 0.0

0.1

0.2

0.3

0.4

0.00 0.0

0.5

0.1

0.2

pga (g)

a)

1.25

1.00

1.00

PFV (m/s)

PFV (m/s)

0.4

0.5

IP_OOP model Transverse direction

1.25

0.75

0.75

0.50

0.50

0.25

0.25

0.00 0.0

0.1

0.2

0.3

0.4

0.00 0.0

0.5

0.1

0.2

pga (g)

0.3

pga (g)

c)

d) 1.0

Longitudinal direction BF IP_OOP

0.8

0.6

Transverse direction BF IP_OOP

0.6

PFV (m/s)

PFV (m/s)

0.5

1.50

IP_OOP model Longitudinal direction

0.8

0.4

b)

1.50

1.0

0.3

pga (g)

0.4

0.2

0.4

0.2

0.0

0.0 0.1

0.2

0.3

pga (g)

e)

0.4

0.5

0.1

0.2

0.3

0.4

0.5

pga (g)

f)

Fig. 7.31 Global results: PFV a BF model (Longitudinal); b BF model (transverse); c IP_OOP model (Longitudinal); d IP_OOP model (transverse); e Comparison (longitudinal) and f Comparison (transverse)

This sub-section is divided into three parts, namely the first part related to the analysis of the peak infill accelerations (PIA) where are presented the range of values reached by the panels, the ratio between the PIA and the pga, the PIA heightwise profiles and finally the infill acceleration response spectra (IARS) and the infill acceleration capacity response spectra (IACS). The second part is related to the analysis of the peak infill velocities (PIV), where the PIV heightwise profiles and infill velocity response spectra (IVRS) are presented. Finally, the third part is related to the

7.5 Case Study

317

8

8

7

7

6

6 5

BF model

Storey level

Storey level

5

PFV BF,average,0.08g

4

PFV BF,average,0.25g PFV BF,average,0.37g PFV BF,average,0.50g

3

BF model PFV BF,average,0.08g

4

PFV BF,average,0.25g PFV BF,average,0.37g PFV BF,average,0.50g

3

IP_OOP model

IP_OOP model 2

PFV IP_OOP,average,0.08g

2

PFV IP_OOP,average,0.08g PFV IP_OOP,average,0.25g

PFV IP_OOP,average,0.25g

1

PFV IP_OOP,average,0.37g

0.1

0.2

0.3

0.4

PFV (m/s)

a)

PFV IP_OOP,average,0.37g

PFV IP_OOP,average,0.50g

Longitudinal direction

0 0.0

1

0.5

0.6

0.7

PFV IP_OOP,average,0.50g

Transverse direction

0 0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

PFV (m/s)

b)

Fig. 7.32 Global results: inter-storey PFV profile a Longitudinal; b Transverse

assessment of the code provisions regarding the OOP acceleration of non-structural elements and compared with the results obtained in this study. Figure 33a shows the evolution of the PIA with the increase of the pga demands, and it is possible to observe a large variation of results, from 0.05 to 3.2 g. The results show some dispersion, which obviously is related to the non-linear behaviour of the global structure which resulted in the amplification of the PIA. The average curve (black dashed line) shows a roughly linear increase of the PIA from 0.25 to 0.92 g, respectively for a pga demand equal to 0.08 and 0.30 g. After that, some non-linearity is visible which is consistent with the high values of ISDmax observed in sub-Sect. 8.5.5.2. Higher dispersion of the PIA results is visible for pga demands larger than 0.40 g. An average PIA equal to 1.30 g was reached for a pga demand equal to 0.50 g. Figure 7.33b shows that the average ratio PIA/pga reduces very slightly from 1.78 to 1.3, respectively for a pga demand equal to 0.08 and 0.50 g. However, the maximum envelope ranges from 13, for a pga value equal to 0.08 g, to 3.8 for pga equal to 0.4 g and after that a slight increase occurred until 6.5. Concerning the comparison between the PFA and the PIA, shown in Fig. 7.33c, d, it is visible that the PIA is always higher than the PFA, and this difference tends to increase with the increment of the pga demand. Two different PFA’s are plotted, namely the PFA of the bottom floor which supports the infill panel (PFAbot ) and the average value between the top and bottom floors bounding the infill panels (PFAave ). From the results it is possible to observe that the shapes of the average curves are similar, although it seems like higher acceleration values are reached by PIA, from 1.2 to 2.2 times, when compared with PFA. The PIA and PFA heightwise profiles are shown in Fig. 7.34, where it is possible to observe that: (i) larger maximum accelerations values are reached by the infill panels when compared with those from the PFA; (ii) the OOP accelerations of the panels tend to increase with the storey height; (iii) the peaks of the infill acceleration do not correspond to the peaks of the floor acceleration, which can be due to the amplification effect caused by the dynamic properties of the panel; (iv) the panels from the 4th storey in the longitudinal direction and from the 8th floor in the transverse

318 4.0 3.5

7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour 15.0

ACC Minimum envelope Maximum envelope Average

ACC Minimum envelope Maximum envelope Average

12.0

3.0

Ratio PIA/pga

PIA (g)

2.5 2.0 1.5

9.0

6.0

1.0 3.0 0.5 0.0 0.0

0.2

0.4

0.0 0.0

0.6

0.2

0.4

pga (g)

a)

Maximum acceleration (g)

3.5 3.0 2.5 2.0

b) 4.0

Longitudinal direction PIA - Average curve PIA - Maximum envelope PIA - Minimum envelope PFAbot - Average curve

3.5

PFAbot - Maximum envelope PFAbot - Minimum envelope PFAaver - Average curve PFAaver - Maximum envelope PFAaver - Minimum envelope

1.5 PIA=pga

1.0

Maximum acceleration (g)

4.0

0.6

pga (g)

0.5

3.0 2.5 2.0

PIA - Average curve Transverse direction PIA - Maximum envelope PIA - Minimum envelope PFAbot - Average curve PFAbot - Maximum envelope PFAbot - Minimum envelope PFAaver - Average curve PFAaver - Maximum envelope PFAaver - Minimum envelope

1.5 PIA=pga

1.0 0.5

0.0 0.0

0.2

0.4

0.0 0.0

0.6

0.2

pga (g)

0.4

0.6

pga (g)

c)

d)

Fig. 7.33 Infill masonry walls OOP accelerations a PIA; b ratio PIA/pga; c comparison PIA versus PFA (longitudinal); and d comparison PIA versus PFA (Transverse) 8

8

7

7 6

6

5

BF model PIA average,0.08g

4

PIA average,0.25g PIA average,0.37g PIA average,0.50g

3

Storey level

Storey level

5

BF model PIA average,0.08g

4

PIAaverage,0.25g PIA average,0.37g PIA average,0.50g

3

IP_OOP model 2

IP_OOP model

PFA average,0.08g

2

PFA average,0.08g PFA average,0.25g

PFA average,0.25g

1

PFA average,0.37g Longitudinal direction

0 0.00

0.25

0.50

0.75

1.00

1.25

Maximum acceleration Amax (g)

a)

1

PFA average,0.37g

PFA average,0.50g

1.50

1.75

Transverse direction

0 0.00

0.25

0.50

0.75

1.00

1.25

PFA average,0.50g

1.50

1.75

Maximum acceleration Amax (g)

b)

Fig. 7.34 Infill masonry walls OOP accelerations: inter-storey PIA profile a Longitudinal; b Transverse

7.5 Case Study

319

direction reached the highest PIA; and (v) larger differences between the PFA and PIA can be found with the increase of the pga demand. The IARS for the panels of each storey are shown in Fig. 7.35 where, additionally, it were included four vertical lines relative to the OOP natural period of the panels place in the longitudinal direction (Pinfills, OOP,x ), in the transverse direction (Pinfills, OOP,z ) and the natural periods of the first (translational along longitudinal direction, P1st,mode,x ) and the second modes of the structure (translational along transverse direction, P1st,mode,z ), respectively represented by olive, navy, orange and magenta dashed lines. Additionally, it is plotted the average value of all the IARS, black solid line, for both the panels distributed along the longitudinal direction (blue line) and transverse direction (red line). From the results, it can be noticed that the IARS peaks do not match with the periods of the panels for the storeys, 2, 3 and 4. However, the IARS major peak corresponds to the natural period of the panels, which is coherent with the high seismic demand that they are subjected to, and is in agreement with the results observed in Fig. 7.34. The remaining and lower peaks occur for periods close to the natural periods of the structure in all the storeys (except in storey 1 where the seismic demand is much reduced and for low periods). It can be also observed that the IARS of the panels along both directions are similar and do not change with the increase of the building height. Figure 7.36 shows the IACS of the panels distributed along both directions and in each building storey. It is visible that for low acceleration demands, the spectral displacements of the panels vary depending on the building direction both in terms of shape and in terms of peaks. This is directly related to the global seismic behaviour of the structure in each direction since it can lead to higher or lower damage levels in the panels. Similarly to what was studied in Sect. 7.5.6.2, the velocities achieved in the central point of the infill panels were recorded and the PIV are plotted in Fig. 7.37a where it is possible to observe a large dispersion of results, but an almost continuous increase of the panel OOP average velocity with the increase of the pga demand. A roughly linear increase of the PIV is visible until 0.52 m/s, which corresponds to a pga equal to 0.3 g, and afterwards keeping more stable around these values, except when the pga is equal to 0.5 g where there is a slight increase around 3%. Obviously, the variations between the maximum and minimum envelop are significant since the maximum PIV is about 1.4 m/s (2 times higher) and the minimum one is around 0.05 m/s (90% lower). From the comparison between the PIV and PFV (Fig. 7.37b, c), it can be noticed that the PFV is always equal or higher than the PIV, which is curious taking is into account the results obtained by the PIA that are clearly higher than the PFA. Along the longitudinal direction, larger differences are visible (1.33 times higher for pga equal to 0.50 g) instead of what is observed in the transverse direction where the larger differences are reached between 0.10 and 0.37 g, with the largest difference around 15%. Figure 7.38 show the IVRS for each storey and again plotted for the all the panels, along both the longitudinal and the transverse directions, being compared with the natural periods of the panels (1st OOP mode) and the first mode in each direction

320

7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour 3.0

3.0

ACC Average AverageOOP,X AverageOOP,Z

2.5

ACC Average AverageOOP,X AverageOOP,Z

2.5

Pinfills,OOP,X

Pinfills,OOP,X

2.0

Pinfills,OOP,Z

Pinfills,OOP,Z

P1st mode,X

P1st mode,X

2.0

P1st mode,Z

P1st mode,Z

Storey level: 2

Sa (g)

Sa (g)

Storey level: 1

1.5

1.5

1.0

1.0

0.5

0.5

0.0 0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

0.0 0.0

4.0

0.5

1.0

1.5

Period (s)

2.0

2.5

3.0

a) 3.0

ACC Average AverageOOP,X AverageOOP,Z

2.5

ACC Average AverageOOP,X AverageOOP,Z

2.5

Pinfills,OOP,X

Pinfills,OOP,X

Pinfills,OOP,Z

Pinfills,OOP,Z

P1st mode,X

2.0

P1st mode,X

2.0

P1st mode,Z

P1st mode,Z

Storey level: 4

Sa (g)

Sa (g)

Storey level: 3

1.5

1.5

1.0

1.0

0.5

0.5

0.5

1.0

1.5

2.0

2.5

3.0

3.5

0.0 0.0

4.0

0.5

1.0

1.5

Period (s)

2.0

2.5

3.0

c) AverageOOP,Z

2.5

ACC Average AverageOOP,X AverageOOP,Z

2.5

Pinfills,OOP,X

Pinfills,OOP,X

2.0

Pinfills,OOP,Z

Pinfills,OOP,Z

P1st mode,X

P1st mode,X

2.0

P1st mode,Z

P1st mode,Z

Storey level: 6

Sa (g)

Sa (g)

Storey level: 5

1.5

1.5

1.0

1.0

0.5

0.5

1.0

1.5

2.0

2.5

3.0

3.5

0.0 0.0

4.0

0.5

1.0

1.5

2.0

2.5

3.0

e) 3.0

ACC Average AverageOOP,X AverageOOP,Z

2.5

ACC Average AverageOOP,X AverageOOP,Z

2.5

Pinfills,OOP,X

Pinfills,OOP,X

2.0

Pinfills,OOP,Z

Pinfills,OOP,Z

P1st mode,X

P1st mode,X

2.0

P1st mode,Z

P1st mode,Z

Storey level: 8

Sa (g)

Sa (g)

Storey level: 7

1.5

1.5

1.0

1.0

0.5

0.5

1.0

1.5

2.0

Period (s)

g)

4.0

f)

3.0

0.5

3.5

Period (s)

Period (s)

0.0 0.0

4.0

d) 3.0

ACC Average AverageOOP,X

0.5

3.5

Period (s)

3.0

0.0 0.0

4.0

b)

3.0

0.0 0.0

3.5

Period (s)

2.5

3.0

3.5

4.0

0.0 0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

Period (s)

h)

Fig. 7.35 Infill masonry walls OOP accelerations: IARS a Storey 1; b Storey 2; c Storey 3; d Storey 4; e Storey 5; f Storey 6; g Storey 7 and h Storey 8

7.5 Case Study

321

3.0

3.0

ACC Average AverageOOP,X AverageOOP,Z

2.5

ACC Average AverageOOP,X AverageOOP,Z

2.5

Storey level: 2

Storey level: 1

2.0

Sa (g)

Sa (g)

2.0

1.5

1.5

1.0

1.0

0.5

0.5

0.0 0.0

0.2

0.4

0.6

0.8

0.0 0.0

1.0

0.2

0.4

0.6

a) 3.0

ACC Average AverageOOP,X AverageOOP,Z

2.5

ACC Average AverageOOP,X AverageOOP,Z

2.5

Storey level: 3

Storey level: 4

2.0

Sa (g)

Sa (g)

2.0

1.5

1.5

1.0

1.0

0.5

0.5

0.2

0.4

0.6

0.8

0.0 0.0

1.0

0.2

0.4

Sd (m)

0.6

c)

AverageOOP,Z

ACC Average AverageOOP,X AverageOOP,Z

2.5

Storey level: 5

Storey level: 6

2.0

Sa (g)

Sa (g)

2.0

1.5

1.5

1.0

1.0

0.5

0.5

0.2

0.4

0.6

0.8

0.0 0.0

1.0

0.2

0.4

Sd (m)

0.6

0.8

1.0

Sd (m)

e)

f)

3.0

3.0

ACC Average AverageOOP,X AverageOOP,Z

2.5

ACC Average AverageOOP,X AverageOOP,Z

2.5

Storey level: 7

Storey level: 8

2.0

2.0

Sa (g)

Sa (g)

1.0

d) 3.0

ACC Average AverageOOP,X

2.5

1.5

1.5

1.0

1.0

0.5

0.5

0.0 0.0

0.8

Sd (m)

3.0

0.0 0.0

1.0

b)

3.0

0.0 0.0

0.8

Sd (m)

Sd (m)

0.2

0.4

0.6

0.8

1.0

0.0 0.0

0.2

0.4

Sd (m)

g)

0.6

0.8

1.0

Sd (m)

h)

Fig. 7.36 Infill masonry walls OOP accelerations: IACS a Storey 1; b Storey 2; c Storey 3; d Storey 4; e Storey 5; f Storey 6; g Storey 7 and h Storey 8

322

1.3

1.50

ACC Minimum envelope Maximum envelope Average

1.25

Maximum velocity (m/s)

1.5

7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

PIV (m/s)

1.0

0.8

0.5

0.3

0.0 0.0

1.00

PIV - Average curve PIV - Maximum envelope PIV - Minimum envelope PFV - Average curve PFV - Maximum envelope PFV - Minimum envelope Longitudinal direction

0.75

0.50

0.25

0.2

0.4

0.00 0.0

0.6

0.2

0.4

pga (g)

0.6

pga (g)

a)

b) 1.50

Maximum velocity (m/s)

1.25

1.00

PIV - Average curve PIV - Maximum envelope PIV - Minimum envelope PFV - Average curve PFV - Maximum envelope PFV - Minimum envelope Tranverse direction

0.75

0.50

0.25

0.00 0.0

0.2

0.4

0.6

pga (g)

c)

Fig. 7.37 Infill masonry walls OOP velocities: a PIV; b comparison PIV versus PFV (longitudinal); and c comparison PIV versus PFV (Transverse)

of the structure. Regarding the storey 2, 3 and 4 it is possible to observe two peaks around the 0.75 s and 0.25 s, being the first one close to the structure natural period. The shape of the IVRS changes with the increase of the building height in contrast to what was observed in the IARS results. From the storey 6th to the storey 8th it is visible only one peak in the IVRS curves, for values around 0.75 s and for all the panels, while no significant changes are visible concerning the shape of the IVRS curve. The maximum peak was achieved in the storey 8th and 6th with velocity around 1.5 m/s and 1.25 m/s, respectively.

7.5.6.4

Code Recommendations and Safety Assessment

The estimation of the expected PIA is one of the parameters required for the assessment of the OOP seismic vulnerability of infill walls, which should be compared with the expected resistant capacity to perform the safety assessment. It is recognized that this is a novelty topic that is being under investigation, since due to the several number of variables associated to the infill walls (geometry, material and mechanical properties, openings, workmanship), to the non-linearity (in-plane, OOP and combined in-plane-OOP interaction) makes it very hard to perform with low probability of

7.5 Case Study

323

5.0 4.5

VCC Average AverageOOP,X

4.0

Pinfills,OOP,X

AverageOOP,Z Pinfills,OOP,Z

3.5

4.5

VCC Average AverageOOP,X

4.0

Pinfills,OOP,X

P1st mode,Z

Sv (m/s)

2.0

2.0 1.5

1.0

1.0

0.5

0.5 1.0

1.5

2.0

2.5

3.0

3.5

P1st mode,Z

0.0 0.0

4.0

Storey level: 2

2.5

1.5

0.5

P1st mode,X

3.0

Storey level: 1

2.5

0.0 0.0

AverageOOP,Z Pinfills,OOP,Z

3.5

P1st mode,X

3.0

Sv (m/s)

5.0

0.5

1.0

1.5

2.0

a) 4.5

3.5

2.5 2.0

2.0 1.5

1.0

1.0

0.5

0.5 3.0

3.5

0.0 0.0

4.0

Storey level: 4

2.5

1.5

2.5

P1st mode,X P1st mode,Z

3.0

Sv (m/s)

Storey level: 3

2.0

Pinfills,OOP,X Pinfills,OOP,Z

3.5

P1st mode,X P1st mode,Z

3.0

1.5

0.5

1.0

1.5

c)

3.5

4.0

Pinfills,OOP,X Pinfills,OOP,Z

3.5

P1st mode,X P1st mode,Z

Storey level: 5

2.5 2.0

P1st mode,X P1st mode,Z

3.0

Sv (m/s)

3.0

AverageOOP,Z

4.0

Pinfills,OOP,X Pinfills,OOP,Z

3.5

VCC Average AverageOOP,X

4.5

AverageOOP,Z

4.0

Sv (m/s)

3.0

5.0

VCC Average AverageOOP,X

4.5

Storey level: 6

2.5 2.0

1.5

1.5

1.0

1.0 0.5

0.5 0.5

1.0

1.5

2.0

2.5

3.0

3.5

0.0 0.0

4.0

0.5

1.0

1.5

e) 4.5 4.0

4.0

3.5

2.5 2.0

2.0 1.5

1.0

1.0

0.5

0.5 2.5

3.0

3.5

4.0

Storey level: 8

2.5

1.5

2.0

P1st mode,X P1st mode,Z

3.0

Sv (m/s)

Storey level: 7

Period (s)

Pinfills,OOP,X Pinfills,OOP,Z

3.5

P1st mode,X P1st mode,Z

3.0

AverageOOP,Z

4.0

Pinfills,OOP,X Pinfills,OOP,Z

g)

3.5

VCC Average AverageOOP,X

4.5

AverageOOP,Z

1.5

3.0

5.0

VCC Average AverageOOP,X

1.0

2.5

f)

5.0

0.5

2.0

Period (s)

Period (s)

Sv (m/s)

2.5

d)

5.0

0.0 0.0

2.0

Period (s)

Period (s)

0.0 0.0

4.0

AverageOOP,Z

4.0

Pinfills,OOP,X Pinfills,OOP,Z

1.0

3.5

VCC Average AverageOOP,X

4.5

AverageOOP,Z

4.0

Sv (m/s)

5.0

VCC Average AverageOOP,X

0.5

3.0

b)

5.0

0.0 0.0

2.5

Period (s)

Period (s)

0.0 0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

Period (s)

h)

Fig. 7.38 Infill masonry walls OOP velocities: IVRS a Storey 1; b Storey 2; c Storey 3; d Storey 4; e Storey 5; f Storey 6; g Storey 7 and h Storey 8

324

7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

error. This work aims to be a preliminary study of the dynamic behaviour of infills during a seismic event and seeks to be a first step towards the estimation of the amplification factor of both PIA and PIV that could help in the seismic design and assessment of infilled RC structures. Eurocode 8 [4] suggests a verification to the non-structural elements that, in case of collapse, could lead to high risk to the population or integrity of the building structure, which is the case of the infill panels. The clause 4.3.5.2 of Eurocode 8 suggests adopting the Eq. 7.5 for the estimation of the seismic acceleration Sa,EC8 that the panel is subjected to during the seismic event. In this equation α is the relationship between the pga and the gravity acceleration g, S is the soil coefficient, Ta is the fundamental period of the non-structural element which, in this case, will be considered the OOP natural period of the panels in each direction, T1 is the fundamental period of the structure in the relevant direction under analysis, z is the height of the element above the level of the application of the seismic action and H is the total height of the building structure. ( ] ) ) ( ( Ta 2 z) / 1+ 1− = α.S. 3 1 + − 0.5 H T1 [

Sa,EC8

(7.5)

The Sa,EC8 can be compared with the PIA reached in each storey in order to assess the approximation of this proposal for estimating those accelerations (Fig. 7.39a, b). As expected, the EC8 limit increases with the increasing building height, and it is observed that it is not exceeded for the bottom storeys (1, 2 and 3), and is exceeded for the storeys 4, 5 and 6. Regarding the top storeys, again the proposal is quite close to the one observed during the analysis. It should be mentioned that the EC8 limit lead to very conservative and restrictive code demands, but looking at the average limit curve and PIA curve it can be easily seen, that the EC8 limit is exceeded in the intermediary storeys, with a shift of the curve about 0.25–0.5 g which means that the condition is not satisfied. Looking at the damages observed in structures due to earthquakes (Chap. 2), it is visible that most of the OOP collapses occurred in the intermediary storeys, thus agreeing with the results obtained herein. Finally, from the comparison between the SaEC8 and the PIA, shown in Fig. 7.39c, it is possible to verify that 60% of the values do not satisfy the condition (not in the safe side) and that most of those results are related to the transverse direction, where higher levels of ISDmax occurred as well as consequent levels of damage the panels were subjected to.

7.5.6.5

Summary of the Major Conclusions of Parametric Study 2

The parametric study 2 was carried out aiming to assess the infill panels’ influence in the global seismic response of the structure, through the analysis of the average values of floor accelerations and velocities. The floor response spectra were drawn for the design pga according to Eurocode 8 [4] as well as the capacity response spectra are

325

8

8

7

7

6

6

5

5

Storey level

Storey level

7.5 Case Study

4 3

3 2

2 ACC Average EC8

1

Longitudinal direction 0 0.00

4

0.25

0.50

0.75

1.00

1.25

1.50

1.75

2.00

2.25

ACC Average EC8

1

Transverse direction 0 0.00

2.50

0.25

0.50

0.75

1.00

1.25

1.50

1.75

2.00

2.25

2.50

Maximum acceleration Amax (g)

Maximum acceleration Amax (g)

b)

a) 2.5

2.0

Sa,EC8(g)

1.5

1.0

0.5

PIAX 0.0 0.0

PIAZ 0.5

1.0

1.5

2.0

2.5

PIA(g)

c) Fig. 7.39 Comparison with the EC8 provision a inter-storey PIA (longitudinal); b inter-storey PIA profile (transverse) and c Sa,EC8 versus PIA

also obtained and presented according to the Eurocode 8 [4] proposal. The second goal of this parametric study was to study the seismic response of the infill panels themselves. For this purpose, the OOP accelerations and velocities were recorded for all the input ground-motions. The maximum OOP acceleration and velocity values reached by the infill panels at each floor were obtained and derived. The respective infill acceleration and capacity response spectra and infill velocity response spectra were then derived and are also presented. Finally, the in-elevation amplification of those values are evaluated and compared with acceleration demands suggested by Eurocode 8. From the assessment of the infill panels’ influence in the RC building seismic response, the following observations can be drawn, namely: • The infill panels, globally, may induce an increment of the peak floor acceleration between 22 and 66%; • Globally, it is not possible to identify the contribution of the infill panels regarding the relationship between the peak floor acceleration and the respective pga. The main difference is related to the PFA/pga ratio, which is higher in the case of the IP_OOP model;

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7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

• It is not observed the matching of the storeys where it occurred the maximum peak floor acceleration values in the BF and IP_OOP model, which supports the idea that the infill panels can modify the building response; • From the comparison of spectra, it can be seen in some storeys the existence of peaks for lower periods in the case of the IP_OOP models, when compared with those of BF models; • From the results, it can be noticed that the BF model reached higher displacements demands than the IP_OOP, and on the other hand the higher acceleration demands were reached by the IP_OOP model; • It seems that for lower seismic demands, the model without infill panels achieved the highest peak floor velocity, and for medium-high seismic demands, the IP_OOP model achieves the highest PFV. From the study of the infill panels’ seismic behaviour, the following observations can be drawn: • It was visible that the peak infill acceleration was always higher than the peak floor acceleration, and this difference tends to increase with the increment of the pga demand; • The panels OOP accelerations tend to increase with the storey height; • The peaks of the panels OOP acceleration do not correspond to the peaks of the floor accelerations, which can be due to the amplification effect caused by the dynamic properties of the panel; • It can be also observed that the infill acceleration response spectra of the panels do not changed with the increase of the building height. Some differences were detected in terms of peaks and shape when compared with the seismic floor response spectra; • It was visible that for low acceleration demands, the spectral displacements of the panels varied depending on the building direction both in terms of shape and peaks; • The peak velocities of the panels was always higher than the peak floor velocities. The EC8 proposal to evaluate the OOP accelerations that the panels are subjected to seems to underestimate the accelerations for the intermediary storeys. However, in the top and bottom levels of the structure, the EC8 seems to define a safe limit envelope that was not exceeded by the panels’ accelerations.

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327

7.5.7 Seismic Vulnerability Assessment of Infilled RC Structures Subjected to Mainshock-Aftershock Sequence Parametric Study 3 7.5.7.1

Motivation

The RC structures located in seismic-prone regions are not generally exposed to only a single earthquake event, but to a seismic sequence, starting from the foreshocks, mainshock and finally the aftershocks, or even two mainshocks occurring with a reduced time interval between them. The aftershocks are triggered by the mainshock due to both static and dynamic (transient) stress changes that occur during the earthquake progression. The aftershocks can play a very important role as evidenced in recent earthquake events, such for example in the Gorkha (Nepal) earthquake in 2015 [54, 55]. Seventeen days after the mainshock which had a magnitude of 7.8Mw , a powerful aftershock with a magnitude of 7.3Mw occurred in the district of Dolakha, which resulted in around 200 casualties, 2500 injuries and an increment of the previous number of reinforced concrete (RC) buildings damages/collapses. More than 400 aftershocks, with magnitude higher than 4Mw, including another one of 6.6 M, were reported during the subsequent two months. Post-earthquake survey damage assessment reports mentioned that most of the RC structures’ damages and failures occurred due to the aftershock events. Some authors reported that the characteristics of the aftershocks combined with existing damage level of the structures due to the previous mainshock were the main reasons behind the poor structural response. Multiple earthquake events were also reported in L’Aquila (Italy) in 2009 [56], Lorca (Spain) in 2011 [57, 58], Central Italy in 2016 ([59], [60]) and more recently in Mexico in 2017. In all of them, it was consensual that the aftershocks had an important impact in the observed losses. Typically, the seismic assessment of RC structures has been carried out for an initial undamaged state of the entire structure [90]. However, for post-earthquake scenarios this approach is not the most realistic one, since powerful earthquakes are typically followed by aftershocks sequences (with different magnitude, duration, location and other characteristics) and the damage caused by the mainshock has to be considered in the seismic assessment process. This is justified by the fact that the structural strength, stiffness and displacement capacity is reduced with the increment of the structural damages. In addition, it is important to remark that the most of the RC structures located in the Southern European countries were not designed considering the real contribution of the infill masonry walls to the seismic response of the structure, since only infills gravity load was taken into account in the design process. Therefore, seismic vulnerability assessment of structures in post-mainshock scenarios is critical and needs to be properly addressed, since it is not realistic to perform the seismic safety assessment, in a post-earthquake scenario, neglecting the existing damages and assuming an undamaged state.

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7.5.7.2

7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

Objectives and Methodology

The main objective of this parametric study is to present a seismic assessment framework approach for damaged RC structures due to mainshock-aftershock sequence. The fragility curves of the undamaged structures due to prior mainshock are compared to the fragility curves of the undamaged structures obtained in 7.5.5.2. It is again compared the results obtained with the BF and the IP_OOP models in order to evaluate the influence of the presence of the infill walls when mainshock-aftershock sequence is considered. Aiming to assess the seismic vulnerability of damaged structures due to a previous mainshock it was adopted the following protocol: First, the structure is subjected to a mainshock input with certain PGA; afterwards, the mainshock record is used to be the aftershock record, but with an important difference which consists in the application of reduction factors (0.9, 0.8, 0.7, 0.6, 0.5, 0.4, 0.4, 0.3, 0.2 and 0.1) to obtain the aftershock ground-motion and then performing IDA. A set of different groundmotions’ records were therefore used for the mainshock-aftershock sequence. Note that it were also used different records for the mainshock and aftershock respectively, however always respecting the condition that the mainshock (PGAMS ) is always equal or higher than the aftershock one (PGAMS ) PGAMS ≥ PGAAS . For example, a mainshock with ground-motion record number 1 with PGA equal to 0.4 g is considered for the non-linear dynamic analysis. Then, the considered aftershock is again the ground-motion record number 1 (used as the mainshock) but affected by the reduction factors leading to PGAs of 0.36, 0.32, 0.28, 0.24, 0.2, 0.16, 0.12, 0.08 and 0.04 g. Afterwards, the mainshock is incremented (scaled) to a PGA equal to 0.45 g and then the aftershock record sequence is again re-generated. Combinations between different mainshock and aftershock ground-motions are also considered, but always ensuring that the mainshock PGA is higher than the aftershock PGA. A period of 10 s is included between the mainshock and the aftershock groundmotions in order to represent the situation in which the structure comes to rest after the first event, but is not repaired. This procedure was concluded for each mainshockaftershock (MS-AS) combination when the structure was found to reach the collapse during the mainshock or the aftershock. In Fig. 7.40 it is plotted the methodology framework proposed to evaluate the increase of the structural damage due to the aftershocks. From the results of the mainshock impacts on the structure, four groups of damaged buildings were extracted and subjected to an aftershock and the seismic vulnerability was again assessed for the remaining damage state limits. For instance, buildings with slight damage (Group S) were subjected to IDA to evaluate the corresponding vulnerability to exceed the moderate extensive, partial collapse and collapse. The moderate damaged structures (Group M) were evaluated to exceed the extensive, partial collapse and collapse damage states. The extensive damaged buildings (Group E) were analysed to exceed the partial collapse and collapse damage state. Finally, partially collapsed structures (Group PC) were studied to exceed the collapse. The major goal was to compare the probability of a structure to reach a

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329

Fig. 7.40 Seismic assessment methodology adopted for evaluating the influence of aftershocks in the structural damage

certain damage state with different levels of initial damage due to mainshock impacts. More than 350 analyses were carried out within this study. Figure 7.41 illustrates an example of a mainshock-aftershock sequence according to the described method. A mainshock SA1 with PGA equal to 0.054 g is first followed by an aftershock with the same record affected with the reduction factor 1 (SA1 × 1) (Fig. 7.41a). Then, in the second analysis (Fig. 7.41b) the aftershock is affected by the reduction factor 0.6 to a PGA equal to 0.032 g. Afterwards, in the third analysis (Fig. 7.41c) the mainshock SA1 is followed by an aftershock SA1 × 0.3. This procedure was performed for the reduction factors 1;0.9;…;0.2;0.1 but were not presented in Fig. 7.41 to avoid repetitions. In Fig. 7.41d it is presented a mainshock-aftershock sequence composed by SA1 and SA2 (both with the same PGA) respectively. To ensure that the aftershock SA2 has the same pga of the mainshock, scale factor (SF) is applied. Then, the aftershock SA2 , is also subjected to the reduction factors 1 (Fig. 7.41d), 0.6 (Fig. 7.41e) and 0.3 (Fig. 7.41f).

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7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour Mainshock - SA1

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Fig. 7.41 Example of mainshock–aftershock sequences applied to analyse the seismic vulnerability of damaged structures: a SA1 -SA1 × 1; b SA1 -SA1 × 0.6; c a) SA1 -SA1 × 0.3; d SA1 -SA2 xSFx1; e SA1 -SA2 xSFx0.6; and f SA1 -SA2 × SF × 0.3 (SF- scale factor)

7.5 Case Study

331

Mainshock - SA1

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f)

Fig. 7.41 (continued)

7.5.7.3

Numerical Results and Discussion

The present section aims to assess the effect of the mainshock-aftershock impact in the structural response. Complementary comparisons are carried out to evaluate the influence of the infill walls and the impact of the aftershocks in buildings with four different levels of damage induced by a previous mainshock (slight, moderate, extensive and partial collapse). The results are, again, presented and discussed in terms of ISDMAX evolution and fragility curves.

7.5.7.4

Assessment of the MS-AS Impact in the Structural Seismic Response

Based on the ISDmax evolutions, obtained from the non-linear dynamic analyses carried out for each numerical model (Figs. 7.42 and 7.43), it can be observed the impact of the aftershock in the damaged structures. The results, as already mentioned, were grouped according to the initial damage state of the structure after the mainshock (Slightly damaged S; Moderately damaged M; Extensively damaged E; and Partial 5

5

Direction X Model BF

Moderate Extensive Partial Collapse Collapse Group S Group M Group E Group PC

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Fig. 7.42 Damaged Structures: ISDMAX evolution (direction X) a BF model and b IP_OOP model

332 6

7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour 6

Direction Y Model BF

Moderate Extensive Partial Collapse Collapse Group S Group M Group E Group PC

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Fig. 7.43 Damaged Structures: ISDMAX evolution (Direction Y) a BF model and b IP_OOP model

collapse PC) and the ISDMAX is plotted in function of the aftershock PGA (Figs. 7.42 and 7.43). Regarding the ISDMAX evolution the following observations can be drawn: • In all the numerical models higher ISDMAX are obtained in the direction Y of the structure when compared with direction X, similarly to what happened for the undamaged structures; • BF model: similar results of the Group S and M; higher ISDMAX for the direction Y for all the groups; • IP_OOP model: similar trend is observed in group S, larger increase of the ISDMAX were gathered for the groups M and E. Fragility curves were derived for structures with the four different damage states. The fragility curves of the damaged structures are compared with those obtained for the undamaged structures in order to evaluate the impact of the mainshock-aftershock sequence in the structure response. Looking at the response of the slightly damaged structures (Fig. 7.44) it can be observed that all of them reached the moderate damage state for lower PGA demands (above 0.2 g). It can be observed that for an aftershock with a PGA around 0.05 g, 40% of the bare frame RC structures exceeded the moderate limit state, which the undamaged BF model exceeded only in 20% of the cases (Fig. 7.44a). Similarly, the IP_OOP model presented the same trend (see Fig. 7.44b), by showing higher vulnerability than the undamaged situation of the original structure; in this model some of the slight damage models reached the extensive damage state for which it was also possible to plot the corresponding fragility curve. The structures with initial moderate damage state (Fig. 7.45) presented a similar trend, since the extensive and partial collapse damage states were exceeded for lower PGA demands than for the undamaged structure. For example, if it is fixed 50% of probability of exceedance and is performed the comparison between the corresponding PGA, it is possible to find: (a) in the BF model, the extensive damage is reached for PGA 4 times lower and the partial collapse for 3.4 times lower than the

333

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7.5 Case Study

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Fig. 7.44 Damaged Structures: Fragility curves for slightly damaged structures a BF model and b IP_OOP model

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Fig. 7.45 Damaged Structures: Fragility curves for moderately damaged structures a BF model and b IP_OOP model

undamaged structure; and b) IP_OOP model exceeded the extensive limit state for PGA demand 2 times lower and the partial collapse for a PGA 1.9 times lower. Regarding the structures with initial extensive damage state (Fig. 7.46), larger differences can be observed between the fragility curves of the undamaged and damaged structures, for partial collapse and collapse. Starting from the analysis of the undamaged BF model response, 40% of the probability of exceedance is reached for the partial collapse and collapse states, for the seismic demand of 0.1 and 0.2 g, respectively. On the other hand, the undamaged BF model reached the same probability of exceedance for a pga of 0.62 g and 1 g, respectively (see Fig. 7.46a). This means that the same probability of occurrence the partial collapse and collapse of the damaged structures is reached for a seismic demand 84% and 80% lower, when compared with undamaged ones. Regarding the IP_OOP damaged models response, the 40% of probability of occurrence occurs for the partial collapse and collapse for

334

7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour 1.0

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Fig. 7.46 Damaged Structures: Fragility curves for extensively damaged structures a BF model and b IP_OOP model

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a pga of 0.13 g and 0.22 g, respectively, which are 78 and 67% lower than that of undamaged structures (see Fig. 7.46b). Finally, the vulnerability of the partially collapsed structures is shown in Fig. 7.47. From the results, it can be observed that for a probability of exceedance of 40% occurs for a pga of 0.1 g and 1 g, respectively damaged and undamaged BF model (Fig. 7.47a). The same probability of exceedance (40%), occurs for a pga of 0.18 g and 0.66 g, respectively damaged and undamaged IP_OOP model (see Fig. 7.47b).

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Fig. 7.47 Damaged Structures: Fragility curves for partially collapsed structures a BF model and b IP_OOP model

7.5 Case Study

7.5.7.5

335

Assessment of the Impact of the Infill Masonry Walls in Damaged Structures

In the sequence of the mainshock-aftershock damage assessment it is also important to discuss the influence of the infill walls impact in the structure response and in particular the differences between the models which considered and not the infill walls OOP behaviour. For this, the ISDMAX evolution for each damaged group was plotted in Fig. 7.48. The results were grouped for each direction and it was performed the comparison between the BF and the IP_OOP models response (see Fig. 7.48). For the slightly damaged structures, it can be observed that: (i) 60% of the models did not reached the moderate damage state for a seismic demand lower than 0.2 g; ii) concerning the seismic demands higher than 0.2 g, it can be observed that 70% of the models reached moderate damage state and the remaining 30% extensive damage. Besides the small number of IP_OOP models with prior slight damage, it can be observed that the moderate damage state was reached in some cases for a seismic demand of 0.04 g. Concerning the results of the structures with prior moderate damage state, it can be observed that the partial collapse occurred for both BF and IP_OOP models, for a seismic demand between 0.35 and 0.55 g. The extensive damage is reached by the IP_OOP model for lower seismic demand when compared the BF models. From the observation of the results of the structures with prior extensive damage, it can be observed that in the longitudinal direction the partial collapse occurred for larger seismic demand when compared with the transverse direction. It is also observed that, in the transverse direction, the partial collapse of the BF models occurred for larger seismic demand when compared with the IP_OOP models. Finally, the collapse damage state of the structures with initial partial collapse state, it is visible that it occurred for lower seismic demand in the case of the IP_OOP models when compared with the BF response. From the moderate damage state fragility curve with initial slight damage (Fig. 7.49) it can be observed that the BF models are the most vulnerable ones. They exceeded 50% the probability of occurrence moderate damage for pga values of 0.06 g, which is 50% lower than the observed for the IP_OOP model, respectively. Similarities can be found in the moderately damaged structures fragility curves for extensive damage and partial collapse (Fig. 7.50), since the BF model is the most vulnerable one. However, some differences can be pointed out, namely that at 50% of probability of exceedance the extensive damage state occurred for 0.1 g and 0.18 g by the BF and IP_OOP, respectively. Larger differences can be observed for the partial collapse fragility curve (Fig. 7.50b) where 50% of probability of exceedance occurred for 0.21 g and 0.31 g by the BF and IP_OOP models, respectively. Regarding the extensively damaged structures’ fragility curves, for partial collapse and for collapse damage states (Fig. 7.51), some interesting findings can be addressed: – The IP_OOP model has a lower probability of exceed the partial collapse damage state than that of BF model, for the same pga demand. For example, 50% of probability of exceedance occurred for a pga demand of 0.11 g and 0.17 g, respectively

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7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

Damage after MS

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Fig. 7.48 Damaged Structures: Assessment of the infill walls impact ISDMAX according to each damage group due to mainshock impact

7.5 Case Study

337 1.0

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Fig. 7.49 Damaged structures: assessment of the infill walls impact Fragility curves of slight damaged structures due to mainshock impact

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Fig. 7.50 Damaged structures: assessment of the infill walls impact Fragility curves of moderately damaged structures due to mainshock impact for a Extensive damage; and b Partial Collapse

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Fig. 7.51 Damaged structures: assessment of the infill walls impact Fragility curves of extensively damaged structures due to mainshock impact for: a partial collapse and b collapse

7 Simplified Macro-modelling of Infill Masonry Walls Seismic Behaviour

Fig. 7.52 Damaged structures: assessment of the infill walls impact Fragility curves of partially collapsed structures due to mainshock impact for collapse damage state

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for BF and IP_OOP model. For a pga of 0.2 g, the probability of occurrence the partial collapse is 85 and 60% of the BF and IP_OOP models, as shown in Fig. 7.51a; – Concerning the probability of the structures with initial extensive damage to reach the collapse damage state, is observed again that is higher for the BF models when compared with the IP_OOP models. For example, the probability of occurrence of 50% occurs for a seismic demand of 0.21 g and 0.28 g, respectively BF and IP_OOP models; – Regarding the collapse fragility curve (Fig. 7.51b), it is also observed that the IP_OOP model exceeded the BF curve for PGA higher than 0.31 g. The fragility curves for the structures with initial damage state corresponding to partially collapsed structures (Fig. 7.52) reinforce the trend of the previous results, namely, the fact that the IP_OOP model fragilkity is much closer to the BF model response. In this plot, it can be observed that for aftershocks with PGA larger than 0.32 g the probability of collapse is similar for the BF and IP_OOP structures. The 50% of probability of exceedance of the collapse damage state occurred for a PGA of 0.11 g for the BF structures, which is 0.64 times lower than the one observed for the IP_OOP model.

7.5.7.6

Summary of the Major Conclusions of Parametric Study 3

The main objective of this parametric study was to present a seismic assessment framework approach for damaged RC structures due to mainshock-aftershock sequence. The methodology was deeply described and discussed. Then, fragility curves of the undamaged structures due to prior mainshock were compared to the fragility curves of the undamaged structures. The major goals of this study were: (i) assess the seismic vulnerability of damaged structures; and (ii) study the influence of the infill walls on the seismic response of damaged structures.

7.6 Final Considerations

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From the seismic vulnerability assessment of damaged structures, the following conclusions can be addressed: • Higher seismic vulnerability of the damaged structures when subjected to aftershocks when compared to the undamaged structures response; • Higher seismic vulnerability of the model without infill walls under the mainshock-aftershock sequences, for the four different damaged states analysed; • The infill walls did not present an important influence in the seismic response of damaged structures. This can be justified with their damage in the mainshock event, which reduced the infills strength and stiffness contribution/participation.

7.6 Final Considerations The present chapter was divided into three major parts, namely: (i) state-of-art review concerning numerical strategies to simulate the infill panels’ seismic behaviour; (ii) development, implementation and calibration of a simplified macro-model; and (iii) parametric studies to assess the effect of the infill walls in the seismic response of RC buildings. A simplified macro-model was developed and implemented in OpenSees [3] to simulate the IP and OOP behaviour of infill walls when subjected to seismic loadings. This model is adapted from the typical bi-diagonal strut model which considers the interaction between the infill and the surrounding RC frames. The OOP behaviour was introduced in the numerical model by allocating the infill mass in the two central nodes. This model, with OOP elastic behaviour when subjected to OOP demands, increases the corresponding OOP displacements which are always interacting with the IP drift demands during the seismic analysis. The numerical model was developed to couple the IP and OOP behaviour in the response of the infill walls during an earthquake with a simplified approach. The IP and OOP interaction drift limits were selected based on previous experimental works conducted by other authors. When the numerical model reaches the drift limits, the infill walls are removed through an element removal algorithm. Regarding the exploratory applications that were carried out in the present study, three RC buildings with the same geometric and mechanical properties were subjected to incremental dynamic analysis for three different situations: bare frame, building with infills with only IP behaviour, and building with infills with IP and OOP behaviour. From the effect of the infill panels, and in particular considering the OOP behaviour in the seismic assessment of RC structures the following observations can be drawn: • The infills OOP behaviour may increase the building vulnerability, leading to the collapse of the most vulnerable storeys for peak ground accelerations greater than 0.3 g;

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• For the slight, light and moderate drifts, the BF model was found the most vulnerable one, while for the other drift levels it was the IP_OOP model (even with a regular distribution of infills); • A significant difference was observed between the IP and IP_OOP numerical models, from which it can be concluded the importance of considering the OOP behaviour of IM walls in the seismic safety assessment of existing RC infilled structures; From the assessment of the infill panels’ influence in the RC structure seismic response, the following conclusions can be drawn: • For pga larger than 0.30 g the infills seem not to be protective to the structure and the ISDmax increase quickly due to the progressive levels of seismic demands they are subjected to which cause high levels of damage or even their collapse; • The infill panels, globally, may induce an increment of the peak floor acceleration between 22 and 66%; • Globally, it is not possible to identify the contribution of the infill panels regarding the relationship between the peak floor acceleration and the respective pga. The main difference is related to the PFA/pga ratio, which is higher in the case of the IP_OOP model; • It is not observed the matching of the storeys where it occurred the maximum peak floor acceleration values in the BF and IP_OOP model, which supports the idea that the infill panels can modify the building response; • From the comparison of spectra, it can be seen in some storeys the existence of peaks for lower periods in the case of the IP_OOP models, when compared with those of BF models; • From the results, it can be noticed that the BF model reached higher displacements demands than the IP_OOP, and on the other hand the higher acceleration demands were reached by the IP_OOP model; • It seems that for lower seismic demands, the model without infill panels achieved the highest peak floor velocity, and for medium-high seismic demands, the IP_OOP model achieves the highest PFV. From the assessment of the infill panels’ seismic behaviour, the following conclusions can be drawn: • The peak infill acceleration was always higher than the peak floor acceleration and tend to increase with the increment of the seismic demand; • The panels OOP accelerations tend to increase with the storey height; • It can be also observed that the infill acceleration response spectra of the panels do not changed with the increase of the building height; • For low acceleration demands, the spectral displacements of the panels varied depending on the building direction both in terms of shape and peaks; • The peak velocities of the panels was always higher than the peak floor velocities; The EC8 proposal to evaluate the OOP accelerations that the panels are subjected to, seems to underestimate the accelerations for the intermediary storeys. However

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in the top and bottom levels of the structure, the EC8 seems to define a safe limit envelope that was not exceeded by the panels’ accelerations. From the seismic assessment of damaged structures due to mainshock-aftershock sequences, the following conclusions can be drawn: • Higher seismic vulnerability of the damaged structures when subjected to aftershocks when compared to the undamaged structures response; • Higher seismic vulnerability of the model without infill walls under the mainshock-aftershock sequences, for the four different damaged states analysed; • The infill walls did not present an important influence in the seismic response of damaged structures. This can be justified with their damage in the mainshock event, which reduced the infills strength and stiffness contribution/participation. Finally, it is important to emphasize the role played by the infill walls in the seismic performance of the structures. It can be observed that the response of the bare frame and the infilled structures are significantly different, since the presence of the infills modified the structural response of the building. Not considering the infill walls OOP behaviour can lead to very different predictions of the expected seismic performance of the structure, which, based on the results of the present study, can be considered as a non-conservative strategy. However, it is recognized the performed study needs to be further developed for a larger family of buildings and with more records to allow drawing more solid conclusions; also the study extension to include infill walls with different horizontal and vertical disposition (irregular distribution) is a important step to be done.

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11. Hendry AW (1973) The lateral strength of unreinforced brickwork. The Struct Eng 51(2):43–50 12. Tassios T (1984) Masonry infill and RC walls. In: Third international symposium on wall structures, Warsaw, Poland 13. Thiruvengadam V (1985) On the natural frequencies of infilled frames. J Earthq Eng Struct Dyn 13(3):401–419 14. Chrysostomou C (1991) Effects of degrading infill walls on the nonlinear seismic response of two-dimensional steel frames. Cornell University, PhD Degree 15. Crisafulli F (1997) Seismic behavior of reinforced concrete structures with masonry infills. PhD Thesis, University of Canterbury 16. SeismoSoft (2004). SeismoStruct—a computer program for static and dynamic nonlinear analysis of framed structures. http://www.seismosoft.com 17. Smyrou E, Blandon C, Antoniou S, Pinho R, Crisafulli F (2011) Implementation and verification of a masonry panel model for nonlinear dynamic analysis of infilled RC frames. Bull Earthq Eng 9:1519–1534 18. Skafida S, Koutas L, Bousias SN (2014) Analytical modeling of masonry infilled RC frames and verification with experimental data. J Struct 2014 19. Dakhakhni W, Elgaaly M, Hamid A (2003) Three-strut model for concrete masonry-infilled steel frames. ASCE J Struct Eng 129(2):177–185 20. ANSYS (2012) ANSYS structural analysis software 21. Furtado A, Costa C, Arêde A, Rodrigues H (2016) Geometric characterisation of Portuguese RC buildings with masonry infill walls. Eur J Environ Civil Eng 1–16 22. Trapani FD, Shing PB, Cavaleri L (2018) Macroelement model for in-plane and out-of-plane responses of masonry infills in frame structures. J Struct Eng 144(2):04017198 23. Ricci P, Di Domenico M, Verderame GM (2018) Empirical-based out-of-plane URM infill wall model accounting for the interaction with in-plane demand. Earthq Eng Struct Dyn 47(3):802– 827 24. Mazza F (2019) In-plane–out-of-plane non-linear model of masonry infills in the seismic analysis of r.c.-framed buildings. Earthq Eng Struct Dyn 48(4):432–453 25. Al Hanoun M, Abrahamczyk L, Schwarz J (2019) Macromodeling of in- and out-of-plane behavior of unreinforced masonry infill walls. Bull Earthq Eng 17(1):519–535 26. SAP2000 (2005) Linear and nonlinear static and dynamic analysis and design of threedimensional structures, vol 10. Computers and Structures, Inc., Berkeley 27. Rots J (1991) Smeared and discrete representations of localized fracture. Int J Fract 51(1):45–59 28. Lourenço P (1996) Computational strategies for masonry structures. PhD Thesis, Delft University 29. Mohyeddin A, Goldsworthy HM, Gad EF (2013) FE modelling of RC frames with masonry infill panels under in-plane and out-of-plane loading. Eng Struct 51:73–87 30. Goodman R, Taylor R, Brekke T (1968) A model for the mechanics of jointed rock. ASCE J Soil Mech Found Div 94(3):637–659 31. Kost G, Weaver W, Barber R (1974) Nonlinear dynamic analysis of frames with filler panels. ASCE J Struct Div 100(4):743–757 32. Ahmed H (2017) Seismic risk assessment of reinforced concrete frames with masonry infill. PhD Thesis, University of Porto 33. Rodrigues H, Varum H, Costa A (2010) Simplified macro-model for infill masonry panels. J Earthq Eng 14(3):390–416 34. Crisafulli F, Carr A, Park R (2000) Analytical modelling of infilled frames-structures—a general review. Bull N Z Soc Earthq Eng 33(1):30–47 35. Furtado A, Rodrigues H, Arêde A (2014) Numerical modelling of masonry infill walls participation in the seismic behavior of RC buildings. OpenSees Days Portugal 2014—OPD2014. Portugal 36. Manzouri T (1995) Nonlinear finite element analysis and experimental evaluation of retrofitting techniques for unreinforced masonry structures. PhD thesis, University of Colorado-Boulder 37. Shing P, Stavidris A, Koutromanos I, William K, Blackard B, Kyriakides M, Billington S, Arnold S (2009) Seismic performance of non-ductile RC frames with brick infill. Improving te Seismic Performance of existing buildings and other structures. ATC/S EI

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Chapter 8

Final Conclusions and Future Works

8.1 Conclusions and Major Contributions The work developed in this thesis mainly consisted in studying the seismic behaviour of infill panels, focusing in their IP and OOP behaviour and in their interaction with the RC building structure, and studying of strengthening strategies to improve the infills seismic performance. In order to achieve these goals, five major objectives were defined for this work, namely: (i) material, mechanical and dynamic characterization of panels built with hollow clay horizontal bricks; (ii) experimental assessment of the OOP behaviour of full-scale infill walls with and without previous damage; (iii) evaluation of the efficiency of strengthening solutions to improve the infill walls seismic behaviour and prevent their collapse; (iv) development and implementation of a simplified macro-model accounting with their IP-OOP behaviour interaction in the software OpenSees; and (v) assessment of the infill walls effect in the seismic response of RC buildings. This thesis started with the overview of the importance of the infill walls in the recent seismic events over the world. From that, it was possible to observe that the infill panels, when distributed uniformly in-plane and in-elevation can have a positive contribution to the global response of the building. However, irregularities due to their distribution can increase the potential development of global failure mechanisms such as soft-storey or torsion. The OOP collapse was the most critical issue in recent earthquakes, since it can result in serious human and material consequences. Some authors pointed out that the predicted costs related to the infills, on average, are equal to 50% of the repair costs of the buildings, highlighting the paramount importance of the infill walls in the seismic loss assessment of RC buildings. Then, it is presented a systematic review of the OOP tests available in the literature to highlight the major issues concerning the infill panels OOP behaviour. The complexity inherent to the OOP behaviour of these elements is reflected by the number of variables considered throughout the experimental studies available in the

© The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 A. F. Furtado, Seismic Vulnerability Assessment and Retrofitting Strategies for Masonry Infilled Frame Building, Springer Theses, https://doi.org/10.1007/978-3-031-20372-5_8

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literature, such as the panel geometries, masonry units, openings, boundary conditions, panel support conditions, workmanship, gravity load and previous IP damage. The review provided new findings and correlations that support the characterization of the OOP seismic behaviour of as-built infill panels. It is also clear the high variability of the results found, due to several reasons such as material variability, workmanship, different loading protocols. From the systematic review, the major observations that can be drawn are: – The OOP strength increases with the reduction of the panel slenderness. It was also observed that the OOP strength capacity decreases for panels with larger aspect ratio. The aspect ratio of the panel changes also the cracking pattern and the failure mechanism. The OOP strength increases for panels with higher compressive and flexural strength; – The IP-OOP behaviour interaction is still an open issue due to the reduced amount of data. Nonetheless, it was observed that the previous damage due to IP tests reduce the panel OOP initial stiffness, cracking strength, maximum strength and secant stiffness. It was also observed that could potentiate fragile failures since the IP demands detach the panel from the envelope frame. From the literature review of the available strengthening strategies for infill panels it was found that, two different approaches can be assumed, namely the disconnection of the panel from the superstructure or the effective strengthening of the panel. Different techniques were found concerning the effective strengthening of the infills, all of them with interesting results such in terms of strength, deformation and energy dissipation. All of the strengthening techniques improve the walls behaviour if a proper anchorage of the panel is ensured. No guidelines or simplified procedures were developed or published until the moment concerning the design of strengthening solutions which difficult the implementation of proper measures by the structural designers. An extensive experimental campaign was carried out comprising material and mechanical characterization tests of infill wallets made with hollow clay horizontal bricks. From the tests it was observed a fragile behaviour from this type of infills. Namely, the collapse of parts of bricks with the increment of the vertical compressive loads, lead to higher OOP instability of the panel which can result in fragile collapses of the panels. It was also observed that the plaster contributed to the increase of the infill panels’ flexural strength. Valuable information concerning material and mechanical properties of the infill panels with hollow clay horizontal bricks was provided. Ambient vibration tests were carried out in laboratory and in-situ to obtain the natural frequencies of the infill panels. From that, it was observed that single-leaf panels’ first OOP frequency ranges from 15 to 36 Hz. The openings can reduce the OOP frequencies from 20−40% depending on the typology and the opening area. It was observed that after the fourth day of the panel construction there is no significant variations of their IP and OOP natural frequencies. The experimental characterization of the OOP behaviour of full-scale infill panels was carried out, first resorting to quasi-static OOP tests with airbags and later with

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pneumatic jacks. A total of nine tests were carried out, in which the variables under study were the effect of the columns axial load, effect of reduction of the panel support width, effect of previous damage, effect of plaster, effect of the test setup and effect of the workmanship. From the experimental campaign, the most relevant findings are: – The variables that affected most were the existence of previous damage and the reduction of the panel support width. It was observed that the OOP strength of a panel, first subjected to an IP test with a maximum drift of 0.5%, reduced 73% and their energy dissipation capacity about 91%. The reduction of the panel support width resulted in the reduction of the panel OOP strength and energy dissipation capacity about 59 and 70%, respectively; – Important variations were found due to the test setup in terms of maximum strength and energy dissipation capacity. It was also observed that the plaster contributed to the spread of the panel cracking. The variations due to the workmanship can range from 10 to 30% in all the response parameters. However, Additional tests are needed to clarify the effect of each variable. he efficiency of TRM based solutions to improve the infill panels’ seismic behaviour and prevent their collapse was assessed through the flexural strength tests and full-scale quasi-static OOP tests. From the flexural strength tests, it was observed that the TRM revealed to be effective in the improvement of the flexural capacity of the panels, both in terms strength and deformation. From the OOP tests, it was also found that the TRM solution was efficient since it prevented the collapse of the panel and improved the maximum strength, deformation and energy dissipation capacity of the panels. It was observed that the use of steel connectors for the mesh-frame anchorage was very effective and that the use of steel plate between the frame and the mesh plus the steel connector reduced the sliding failure of the mesh in the regions surrounding the steel connector. The use of a stronger mesh did not resulted in better performance in terms of strength when compared with response of a panel strengthened with a mesh with lower strength characteristics. The strengthening contributed to the increase of the maximum strength between 1.75−2.44 times, until 2.72 times the deformation capacity, and the energy dissipation at least 2 times. The last contribution of this thesis was the development of a simplified macromodel that was later implemented in the software OpenSees. The model is an equivalent double-strut model which consider the interaction between the panel and the frame. The model simulates the IP-OOP behaviour interaction within an ongoing simulation and allows to define an IP-OOP drift limits in which when the curve limit is exceeded the panel is removed. Several non-linear dynamic analyses were carried out from which the following conclusions were drawn: – The infill panels have can have a positive contribution to the seismic response of the structure for lower seismic demands. Different responses were found considering and not he IP-OOP behaviour interaction. The IP-OOP interaction increased the vulnerability of the structure for medium-high seismic demands. The infills OOP behaviour may increase the building vulnerability, leading to the collapse of the

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8 Final Conclusions and Future Works

most vulnerable storeys for peak ground accelerations greater than 0.3 g. This fact highlight the importance of considering the OOP behaviour of the infill walls; The infill panels, globally, may induce an increment of the peak floor acceleration from 22−66%. Globally, it is not possible to identify the contribution of the infill panels regarding the relationship between the peak floor acceleration and the respective pga. It seems that for lower seismic demands, the model without infill panels achieved the highest peak floor velocity, and for medium-high seismic demands, the IP_OOP model achieves the highest PFV; The peak infill acceleration was always higher than the peak floor acceleration and tend to increase with the increment of the seismic demand. The panels OOP accelerations tend to increase with the storey height. The peak velocities of the panels was always higher than the peak floor velocities; The EC8 proposal to evaluate the OOP accelerations that the panels are subjected to, seems to underestimate the accelerations for the intermediary storeys. However, in the top and bottom levels of the structure, the EC8 seems to define a safe limit envelope that was not exceeded by the panels’ accelerations; Higher seismic vulnerability of the damaged structures when subjected to aftershocks when compared to the undamaged structures response; Higher seismic vulnerability of the model without infill walls under the mainshock-aftershock sequences, for the four different damaged states analysed; The infill walls did not present an important influence in the seismic response of damaged structures. This can be justified with their damage in the mainshock event, which reduced the infills strength and stiffness contribution/participation.

Finally, it is important to emphasize the importance of the infill walls in the seismic performance of the structures. It can be observed that the response of the bare frame and the infilled structures are significantly different, since the presence of the infills modified the structural response of the building. Not considering the infill walls OOP behaviour can lead to very different predictions of the expected seismic performance of the structure, which, based on the results of the present study, can be considered as a non-conservative strategy.

8.2 Future Works The work presented in this thesis contains useful contributions for the characterization of the mechanical and material properties of infill panels, assessment of the OOP vulnerability and development and evaluation of strengthening strategies, and numerical modelling and assessing of infilled RC structures. Nevertheless, additional research would be important to consolidate the knowledge acquired and to reach other capabilities, making the findings and procedures used and tested in this work even more useful and common, both in the scientific community and in the industry. In this sense, a few research topics related with the developed work are presented in the following paragraphs:

8.2 Future Works

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– Systematic review of the IP tests available in the literature (collect detailed information from each specimen in terms of geometric dimensions, material and mechanical properties, test setup and instrumentation details, and the discussion of major results and findings; – Development and validation of analytical prediction models to estimate the OOP strength capacity of infill panels, with and without previous damage, according to code standards, models available in the literature, cracking line theory; – Additional experimental tests comprising compression strength tests and diagonal tensile strength tests of infill panels built with horizontal hollow clay bricks with plaster. Repetition of the flexural strength tests, parallel and perpendicular to the horizontal bed joints, in panels with higher geometric dimensions (height/width ratio) in panels with and without plaster. The results will be useful to validate and results herein presented. Initial shear strength and bond strength tests would be very interesting to perform and complete the characterization of the infill properties. All those characterization tests should be carried out in panels built with different masonry units, such as vertical concrete hollow blocks, vertical clary hollow bricks, very common nowadays in the Portuguese construction; – Additional ambient vibration tests should be performed in panel with similar geometric properties of the specimen tested in Chap. 5 to compare the results among them and validate the findings herein presented. Tests in panels with different typologies, geometric dimensions and boundary conditions should increase the available results where it could be possible to extract solid conclusions regarding the influence of each variable in the OOP vibration modes of the panel. In the same way, the replication of the tests in panels built with different masonry units are also needed; – Better characterization of the uncertainties associated to the ambient vibration tests are also needed. The optimization of the methodology herein presented, could also explore new developments concerning the damage detection in infill panels in after-earthquake scenarios. For this, appropriate analytical tools should be developed to calculate the OOP vibration modes of infill panels with different characteristics in its as-built condition. Then, those expected frequencies are compared with the ambient vibration tests carried out in their damage state, and based on their higher modes, it will be possible to localize the damage in the panel; – Perform additional OOP tests of full-scale infill panels assessing the influence of parameters such as slenderness (panels with hollow clay horizontal bricks with different thicknesses), boundary conditions (panels built with only three or two edges mortared); different levels of previous damage due to prior IP test (0.1; 0.2, …, 0.9 and 1%); panels built with the same geometry and material properties to assess the variability among the tests, panels built with three types of workmanship to assess the variations among them; panels with reduced width support with previous damage. The assessment of the IP response of the infilled RC frame of panels prior subjected to different levels of OOP damage; – Perform tests comprising combined IP-OOP loadings with an improved test setup would help to assess the infill panels OOP capacity with different levels of distortion (IP drift) of the envelope frame;

352

8 Final Conclusions and Future Works

– Material characterization tests of different textile reinforced mortar solutions, by carrying out tensile strength tests of dog bones specimens; – Optimization of the number of steel connectors and the dimension of the connector. It should be also taking into account that the definition of the position of the metallic connectors must consider the columns and beams coating to avoid their application along the steel bar alignment; – The definition of the plastic connectors for the panel-mesh anchorage must be defined according to the bed-joints alignments, since the anchorage is more effective when the drill were effectuated in the bed joints. It provide higher tensile strength to the mesh and robustness to the connector fixation; – Optimization of the mesh since the major aspect is the anchorage of the mesh to the frame, the mesh can be adjusted to one with lower strength properties and more ductility; – Cost-benefit analysis of the strengthening solutions available and tested; – Calibration of the out-of-plane behaviour in the simplified numerical model and the calibration of the IP-OOP displacement interaction law based on experimental evidences, for panels built with different types of masonry units; – Assessment of the infill panels effect in the seismic response of RC buildings with different number of storeys.

Author Biography

André Furtado concluded his PhD in Civil Engineering, with specialization in engineering structures, at FEUP in 2020. He has experience in seismic and structural engineering, numerical modelling and experimental testing of RC buildings. Currently, He is Assistant Professor in Structural Engineering at the Instituto Superior Técnico. He has participated actively in 6 national and 4 international research projects. Also, He currently participates in 3 Cost Actions. He is currently the PI of the pilot project SafEnergy which aims to develop and validate low-cost seismic plus energy techniques for envelopes of RC buildings subjected to pure out-of-plane seismic loadings. He is also Co-PI of the international project ERES2 developed in collaboration with Indian Universities. He is co-author of 1 book, 65 papers in international peer-review journals, 48 of them as the first author. He is also co-author of 10 papers in national peer-review journals, 7 book chapters and 72 papers in international and national conferences. He has won 5 different awards related with his research works. He was the organizer of 1 conference and 2 workshops. In addition, He is/was a member of the organization and scientific committees of national and international conferences. He is a peer-reviewer and member of the editorial board of several peer-review international journals. He is the lead guest editor of 3 special issues. He is currently the co-supervisor of 6 PhD theses and co-supervisor of 20 MSc thesis. Finally, He is also the author of 23 projects/consulting works for the industry community.

© The Editor(s) (if applicable) and The Author(s), under exclusive license to Springer Nature Switzerland AG 2023 A. F. Furtado, Seismic Vulnerability Assessment and Retrofitting Strategies for Masonry Infilled Frame Building, Springer Theses, https://doi.org/10.1007/978-3-031-20372-5

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