Innovative Solutions for Deep Foundations and Retaining Structures: Proceedings of the 3rd GeoMEast International Congress and Exhibition, Egypt 2019 on Sustainable Civil Infrastructures – The Official International Congress of the Soil-Structure Interaction Group in Egypt (SSIGE) [1st ed. 2020] 978-3-030-34189-3, 978-3-030-34190-9

Table of contents :
Front Matter ....Pages i-x
Comparison of Russian Norms (Snip) and European Norms (Eurocodes) for Design and Construction Pile Foundations (Safak Soylemez)....Pages 1-6
Development of Earth-Based Mortars for Usage in Earth Construction (Fady Halim, Rana Khali, Safwan Khedr, Mohamed Darwish, Amani Saleh, Dalia El-Arabi et al.)....Pages 7-17
Experimental Investigations of Capacity Response of Root Piles on Combination of O-Cell Test and Conventional Head-Down Test Methods (Xiao-juan Li, Guo-liang Dai, Wei-ming Gong, Ming-xing Zhu)....Pages 18-39
Finite Element Modeling of Soil Arching in Pile Supported Embankment: 2D Approach (Naveen Kumar Meena, Sanjay Nimbalkar, Behzad Fatahi)....Pages 40-50
Study of Slope Stability Using Piles, Pathological Case in Mostaganem, Algeria (A. Belghit, S. M. A. Bourdim, A. Benanane, N. Bouhamou)....Pages 51-60
Effect of Fines and Matric Suction on the Collapsibility of Sandy Soils (Mohamed A. Alassal, Asmaa M. Hassan, Hussein H. Elmamlouk)....Pages 61-72
Assessment of Earth Retaining Wall Sustainability: Value Functions and Stakeholder Weighting Sensitivity (I. P. Damians, R. J. Bathurst, A. Lloret, A. Josa, D. El Mansouri)....Pages 73-88
Numerical Study of Passive Earth Pressure on Retaining Walls (Meriem Fakhreddine Bouali, Mounir Bouassida)....Pages 89-105
Understanding Pile Foundations Design Through Case Histories of New Tagus Bridge and Leziria Bridge (Pedro S. Sêco e Pinto, Ricardo Oliveira, Alexandre Portugal)....Pages 106-153
Enhancement of Raft Foundation Using Micro Pile Technique (Ahmed T. Farid, Mostafa A. Yousef)....Pages 154-164
Back Matter ....Pages 165-165

Citation preview

Sustainable Civil Infrastructures

Pedro Pinto Chang-Yu Ou Hany Shehata Editors

Innovative Solutions for Deep Foundations and Retaining Structures Proceedings of the 3rd GeoMEast International Congress and Exhibition, Egypt 2019 on Sustainable Civil Infrastructures – The Official International Congress of the Soil-Structure Interaction Group in Egypt (SSIGE)

Sustainable Civil Infrastructures Editor-in-Chief Hany Farouk Shehata, SSIGE, Soil-Interaction Group in Egypt SSIGE, Cairo, Egypt Advisory Editors Khalid M. ElZahaby, Housing and Building National Research Center, Giza, Egypt Dar Hao Chen, Austin, TX, USA

Sustainable Infrastructure impacts our well-being and day-to-day lives. The infrastructures we are building today will shape our lives tomorrow. The complex and diverse nature of the impacts due to weather extremes on transportation and civil infrastructures can be seen in our roadways, bridges, and buildings. Extreme summer temperatures, droughts, flash floods, and rising numbers of freeze-thaw cycles pose challenges for civil infrastructure and can endanger public safety. We constantly hear how civil infrastructures need constant attention, preservation, and upgrading. Such improvements and developments would obviously benefit from our desired book series that provide sustainable engineering materials and designs. The economic impact is huge and much research has been conducted worldwide. The future holds many opportunities, not only for researchers in a given country, but also for the worldwide field engineers who apply and implement these technologies. We believe that no approach can succeed if it does not unite the efforts of various engineering disciplines from all over the world under one umbrella to offer a beacon of modern solutions to the global infrastructure. Experts from the various engineering disciplines around the globe will participate in this series, including: Geotechnical, Geological, Geoscience, Petroleum, Structural, Transportation, Bridge, Infrastructure, Energy, Architectural, Chemical and Materials, and other related Engineering disciplines.

More information about this series at http://www.springer.com/series/15140

Pedro Pinto Chang-Yu Ou Hany Shehata •

•

Editors

Innovative Solutions for Deep Foundations and Retaining Structures Proceedings of the 3rd GeoMEast International Congress and Exhibition, Egypt 2019 on Sustainable Civil Infrastructures – The Official International Congress of the Soil-Structure Interaction Group in Egypt (SSIGE)

123

Editors Pedro Pinto International Society for Soil Mechanics and Geotechnical Engineering (ISSMGE) Portugal, Portugal

Chang-Yu Ou Chinese Taipei Geotechnical Society Taiwan, Taiwan

Hany Shehata Soil-Structure Interaction Group in Egypt (SSIGE) Cairo, Egypt

ISSN 2366-3405 ISSN 2366-3413 (electronic) Sustainable Civil Infrastructures ISBN 978-3-030-34189-3 ISBN 978-3-030-34190-9 (eBook) https://doi.org/10.1007/978-3-030-34190-9 © Springer Nature Switzerland AG 2020 This work is subject to copyright. All rights are reserved 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

Contents

Comparison of Russian Norms (Snip) and European Norms (Eurocodes) for Design and Construction Pile Foundations . . . . . . . . . . Safak Soylemez Development of Earth-Based Mortars for Usage in Earth Construction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Fady Halim, Rana Khali, Safwan Khedr, Mohamed Darwish, Amani Saleh, Dalia El-Arabi, Sara Henry, Ali El-Menoufy, and Joe Said Experimental Investigations of Capacity Response of Root Piles on Combination of O-Cell Test and Conventional Head-Down Test Methods . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Xiao-juan Li, Guo-liang Dai, Wei-ming Gong, and Ming-xing Zhu

1

7

18

Finite Element Modeling of Soil Arching in Pile Supported Embankment: 2D Approach . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Naveen Kumar Meena, Sanjay Nimbalkar, and Behzad Fatahi

40

Study of Slope Stability Using Piles, Pathological Case in Mostaganem, Algeria . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . A. Belghit, S. M. A. Bourdim, A. Benanane, and N. Bouhamou

51

Effect of Fines and Matric Suction on the Collapsibility of Sandy Soils . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Mohamed A. Alassal, Asmaa M. Hassan, and Hussein H. Elmamlouk

61

Assessment of Earth Retaining Wall Sustainability: Value Functions and Stakeholder Weighting Sensitivity . . . . . . . . . . . . . I. P. Damians, R. J. Bathurst, A. Lloret, A. Josa, and D. El Mansouri

73

Numerical Study of Passive Earth Pressure on Retaining Walls . . . . . . Meriem Fakhreddine Bouali and Mounir Bouassida

89

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Contents

Understanding Pile Foundations Design Through Case Histories of New Tagus Bridge and Leziria Bridge . . . . . . . . . . . . . . . . . . . . . . . . 106 Pedro S. Sêco e Pinto, Ricardo Oliveira, and Alexandre Portugal Enhancement of Raft Foundation Using Micro Pile Technique . . . . . . . 154 Ahmed T. Farid and Mostafa A. Yousef Author Index . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 165

About the Editors

Chang-Yu Ou is Professor of Construction Engineering at the National Taiwan University of Science and Technology in Taipei. Professor Ou’s focus of research is on studies of soil behavior and excavation problems, and he has published many journal and conference papers on these subjects. In parallel to his academic career, he has worked closely with industrial builders, taking part in many large-scale excavation projects, gaining valuable practical experience in analysis and design. (1) Chair Professor of Department of Construction Engineering, National Taiwan University of Science and Technology (2) Coordinator of Civil Engineering Program of the National Science Council (3) Editor-in-Chief, Journal of Geoengineering (4) Author of the “Deep Excavation: Theory and Practice” Book

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About the Editors

Hany Shehata is the founder and CEO of the Soil-Structure Interaction Group in Egypt “SSIGE.” He is a partner and Vice-President of EHE-Consulting Group in the Middle East and managing editor of the “Innovative Infrastructure Solutions” journal, published by Springer. He worked in the field of civil engineering early, while studying, with Bechtel Egypt Contracting & PM Company, LLC. His professional experience includes working in culverts, small tunnels, pipe installation, earth reinforcement, soil stabilization, and small bridges. He also has been involved in teaching, research, and consulting. His areas of specialization include static and dynamic soil-structure interactions involving buildings, roads, water structures, retaining walls, earth reinforcement, and bridges, as well as, different disciplines of project management and contract administration. He is the author of an Arabic practical book titled “Practical Solutions for Different Geotechnical Works: The Practical Engineers’ Guidelines.” He is currently working on a new book titled “Soil-Foundation-Superstructure Interaction: Structural Integration.” He is the contributor of more than 50 publications in national and international conferences and journals. He served as a co-chair of the GeoChina 2016 International Conference in Shandong, China. He serves also as a co-chair and secretary general of the GeoMEast 2017 International Conference in Sharm El-Sheikh, Egypt. He received the Outstanding Reviewer of the ASCE for 2016 as selected by the Editorial Board of International Journal of Geomechanics.

Pedro Pinto ISSMGE Past President Education 1965–1971—Licenciated in Civil Engineer (6-year course) (with honors). 1975–1977—Master of Engineering (with honors) 1979–1983—Specialist in Geotechnique (Ph.D. Degree) (with honors) 1992—Director of Research (Full Professor Degree) (with high honors)

About the Editors

ix

Positions – – – – – – – – – – –

ISSMGE Board Member (2017–2021) ISSMGE Immediate Past President (2009–2013) ISSMGE President (2005–2009) ISSMGE Vice President for Europe (2001–2005) Full Professor of Geotechnical Engineering of University of Coimbra. World Bank Consulting for Dam Safety (2013–2015). Invited Professor of Master Courses “Soil Mechanics” and “Engineering Geology” of New University of Lisbon (1983–1995). United Nations Consulting for Design and Instrumentation for Dams (1988–1992). Invited Guest Lecturer of University of California, USA (1992–1994). Chairman of TC4 “Earthquake Geotechnical Engineer” Committee of ISSMGE (1994–2000). President of Portuguese Society for Geotechnique (1996–2000).

Professional Experience Consulting Engineer of 450 major projects in dams, power plants, bridges, tunnels and quay walls, in Portugal, Angola, Argelie, Brazil, Cabo Verde, China, Dominican Republic, Ecuador, Guinea-Bissau, Guinea Ecuatorial, India, Lebanon, Malawi, Morocco, Mozambique, Senegal, Syria, Tunisia, Uganda, Venezuela and Zambia, covering field and laboratory testing, dynamic analyses, earthquake engineering, numerical analyses, ground improvement, slopes, special foundations, instrumentation and safety evaluation. Conferences He has presented more than 350 state-of-the-art Lectures and Special Lectures in 80 countries of the five continents.

x

About the Editors

Awards and Honors He has received more than 50 international awards including American Biographical Institute, USA, “Special Volume for the Contributors of Earthquake Engineering, Nagadi Lecture by Indian Geotechnical Society, Széchy Lecture by Hungarian S M Society and Hungarian Academy of Sciences, Nonveiller Lecture- by Croatia Geotechnical Society, Sukle Lecture by Slovenia Soil Mechanics Society, Chin Lecture by Huanzhou University (China), Qian Jia Huan Lecture by Hohai University (China), and Chin Fung Kee Memorial Lecture by Institute of Engineers of Malaysia. Editorial Boards and Reviewer – Editor of International Journal of Case Histories – Co-editor of Geotechnical and Geological Engineering Journal, Springer Publisher – Member of Editorial Board of several Journals, namely “Geotecnia,” “Bulletin of Earthquake Engineering,” “Acta de Geotecnia,” “International Journal of Geotechnical Engineering.” – Editor of Proceedings of four International Conferences. Publications He is author or co-author of 500 Technical and Scientific Reports, more than 180 papers for national and international conferences and journals and has contributed for ten books.

Comparison of Russian Norms (Snip) and European Norms (Eurocodes) for Design and Construction Pile Foundations Safak Soylemez(&) Enka Ins., Ve San. A.S., Istanbul, Turkey [email protected]

Abstract. In this study, the general principles of pile design and construction according to Eurocode 7 and SP 24.13330.2011 regulations are investigated. Although pile foundations are very common in both Russia and Europe, there are many different understanding that starts from soil investigation stage to construction stage. To evaluate their differences, the minimum requirements for soil investigations, suggested approaches for pile foundation design and construction tolerances are comparatively analyzed. Additionally, bearing capacity of pile foundations calculated from suggested approaches are comparatively evaluated and pile load test results for two real cases located in Iraq and Russia comparatively evaluated. British Standards in compliance with Eurocode 7, requires a more comprehensive site characterization and pile load testing than Russian norms in both preliminary and construction stage. According to the results of analyses, adopting Eurocode 7 and SP 24.13330.2011 codes have similarity in design for ultimate and serviceability limit states. They both adopt partial safety factors to loads and resistances for the ultimate limit state design of the design of pile foundation. However, they visibly differ in terms of the number and content of the design checks, and in terms of the actual values of the partial safety factors. The pile capacities calculated from Russian Norms for both bored and driven is significantly higher than Eurocode 7. Full scale pile loading test results are compared for obtained pile vertical load capacities.

1 Introduction The allowable bearing load determination of bored or driven piles in layered soil strata is a routine engineering task. Although the rules of soil mechanics and general principles of pile design is universal, different codes provide different approaches to determine pile loads. Even though the engineer uses his/her own judgement, knowledge and experience, the design have to be in comply with particular code in force in the relevant juridistinction. In this study, the Eurocode 7 and SP 24.13330.2011 regulations for pile design are investigated. The differences between two code and their effects on the pile design is discussed. In addition to that, the minimum requirements for soil investigations, suggested approaches for pile foundation design and construction tolerances are comparatively analyzed. Additionally, two pile load test results were evaluated to comparison. © Springer Nature Switzerland AG 2020 P. Pinto et al. (Eds.): GeoMEast 2019, SUCI, pp. 1–6, 2020. https://doi.org/10.1007/978-3-030-34190-9_1

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2 Methodology In Eurocode 7 Section 7, the general principles for pile design regardless pile construction methodology (driven, bored, cfa etc.) is given. According to this section, calculations must be performed to show pile capacity is not exceeding under compression, tension load. Additionally, the serviceability condition of the pile has to be checked. Pile load capacity can be calculated by using static and dynamic pile load tests, empirical relationships as well as performance based design based on soil investigation tests. Most important difference, the soil parameters have to be determined via a statistical method in accordance with GOST 20522. Even though Eurocode allows to use statistical method, it is not compulsory to use a statistical method to determine soil parameters. In SP 24.13330.2011 Section 7 gives information for limit state design of piles Section 4 gives an information related to pile serviceability conditions. Russian Code Snip uses both in-situ tests and theoretical calculations given in the Section 7. Eurocode is not restricted the theoretical method used in pile design calculation and allows “engineering judgement”, however Snip dictates a pile design method and correlation coefficient depends on index characteristics of the soil. 2.1

Russian Norms

According to SP 24.13330.2011 ultimate capacity of the pile is determined based on empirical tables and previous experience. The bearing capacity equation is given below. Fd ¼ cc ðccR RA þ uRccf f i hi Þ where – cc service factor of pile in soil, to be assumed equal to 1; R – design resistance of soil bottom end of pile, kPa; A – area of pile support onto the soil, m2, assumed as cross-sectional area of gross pile or according to cross sectional area of springing with respect to its maximum diameter or to the area of net shell-pile.; u – external perimeter of pile shaft cross section, m; fi – design resistance of i soil layer on pile side surface, kPa; hi – thickness of i soil layer which is in contact with pile side surface, m; ccR , ccf – soil service factors under the bottom end and along the side surface of pile respectively. The resistance of both end bearing and skin friction of piles basically depend on liquidity index, fine content and the effective confining pressure. 2.2

European Norms

Eurocode 7, Section 7 of EN 1997-1 for the design of pile foundations can be used for end bearing, friction and tension piles as well as transversely loaded piles. The equilibrium equation to be satisfied in the ultimate limit state design of axially loaded piles in compression is Fc;d  Rc;d. Where Fc;d is the design axial compression load, Rc;d is the pile compressive design resistance.

Comparison of Russian Norms (Snip) and European Norms (Eurocodes)

3

Rc;d ¼ Rc; k/ct Rc;d ¼ Rb;k =cb þ Rs;k =cs EN 1997-1 provides different sets of recommended partial resistance factor values for driven, bored and CFA piles in Tables A6, A7 and A8 of Annex A. Also different design approaches are given in Eurocode. Such as; DA1:C1 : A1 þ M1 þ R1 DA1:C2 : A2 þ M1 or M2 þ R4 DA2:C1 : A1 þ M1 þ R2 DA3:C1 : ðA1 or A2Þ þ M2 þ R3 “A” defines the loads acting on the structures and “M” defines Material Resistance Factors. The allowable pile capacity is determined by using partial factors defined in Table 1. Within this study, Eurocode Design Approach 1 is used. Table 1. Recommended partial resistance factors in Eurocode (compression) Resistance For driven piles Base Shaft Total/Combined For bored piles Base Shaft Total/Combined

Symbol Set R1

R2 R3 R4

cb cs ct

1,0 1,0 1,0

1,1 1,0 1,3 1,1 1,0 1,3 1,1 1,0 1,3

cb cs ct

1,25 1,1 1,0 1,6 1,0 1,1 1,0 1,3 1,15 1,1 1,0 1,5

3 Results and Discussion 3.1

Pile Load Test Results

Mazurkiewicz’s method is a graphical method that is internationally recognized and widely used in engineering practice and applied for all load tests. This method uses the ‘test load vs pile’s settlement’ curve assumed as a parabola drawn up to the maximum test load. A series of equally spaced lines parallel to the load axis are drawn to intersect with the actual load-settlement curve depicted from test data. Then, from each of these intersection points, a line is drawn parallel to the settlement axis, cutting and crossing the load axis. At each intersection point with the load axis, a 45º line is drawn to intersect with the next line parallel to settlement axis. These

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intersections fall, approximately, on a straight line intersecting the load axis at a point which defines the ultimate load. These straight lines were drawn by using least squared method. The ultimate capacities calculated from pile load tests for bored and driven piles in Iraq are given in Figs. 1 and 2.

Fig. 1. Pile load test evaluation for bored pile in Iraq (80 cm diameter)

The allowable capacities of pile calculated in accordance with Eurocode and Snip is given in Table 2 with pile load tests evaluation results. According to the results of analyses, adopting Eurocode 7 and SP 24.13330.2011 codes have similar approaches to determine allowable pile capacities. Both code use the partial factor of safeties to determine allowable pile load capacity. However, this two code visibly differ in terms of the number and content of the design checks, and in terms of the actual values of the partial safety factors. The pile capacities calculated from Eurocode 7 for bored pile is %35 higher than Russian Norms in dominantly cohesive soils. Likely, the calculated pile capacity for driven pile in dominantly cohesive soil %15 percent lower than Eurocode 7. Full scale pile loading test results are compared for obtained pile vertical load capacities. According to the results of field test, both Eurocode and Snip Norms are providing an allowable bearing capacity on the safe side with used factor of safety even though the European Norms gives slightly higher bearing capacity than pile load test results for bored piles. The allowable pile axial bearing capacity obtained from pile load tests are quite similar with calculated allowable capacity with the partial factors from Eurocode 7, however, SNIP Norms

Comparison of Russian Norms (Snip) and European Norms (Eurocodes)

5

Fig. 2. Pile load test evaluation for driven pile in Russia (40 cm * 40 cm) Table 2. Allowable pile capacities Case Russian norms - theoretical European norms - theoretical Pile load test results

Iraq – bored pile (kN) Russia – driven pile (kN) 2010 730 2709 854 2566 884

gives quite conservative results for both driven and bored piles compared with pile load test results.

4 Summary and Conclusions Within this paper, axial pile bearing capacities obtained from field pile loading tests is compared with theoretical calculations based on SP 24.13330.2011 (Updated version of SNIP 2.02.03-85) and Eurocode 7. Normative pile bearing capacity calculated values in accordance with SNIP based on liquidity index, fine content and the effective confining pressure leads to significantly low values when compared with and Pile Load Test results. On the contrary, theoretical calculations with accepted method in literature based on index characteristic of soils shows an agreement with pile load test results.

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A lower factor of safety has been using in SNIP norms, therefore this approach partially cover the calculated ultimate capacity of the piles. Still, the SNIP Norms provides a conservative method whereas Eurocode allows a good agreement with pile load tests results if correct soil parameters are selected in the pile design. Acknowledgments. The author grateful to Enka Insaat ve Sanayi A.S. for their support and permission to use their data.

References BS 8004: Code of practice for foundations. British Standards Institution, London, UK (2015) BS EN 1997-1: Eurocode 7 -1, Geotechnical Design; General Rules. British Standards Institution, London, UK (2010) GOST 5686-94: Soils-Field Test Methods by Piles (1994) GOST 20522-96: Soils-Statistical Treatment of Test Results (1996) Randolph, M.F.: A theoretical study of the performance of piles. Ph.D. thesis, University of Cambridge, UK (1977) Randolph, M.F., Carter, J.P., Wroth, C.P.: Driven piles in clay – the effects of installation and subsequent consolidation. Geotechnique 29(4), 361–393 (1979) SNiP: SNiP 2002.03-85: Pile foundations. Federal Registry of National Building Codes and Standards of Russia, Moscow, Russia (1985) SNiP: SP 24.13330-2011: Pile foundations. Federal Registry of National Building Codes and Standards of Russia, Moscow, Russia (2011) Mazurkiewicz, B.K.: Test loading of piles according to Polish Regulations. Preliminary Report No. 35, Commission on Pile Research, Royal Swedish Academy of Engineering Services, Stockholm (1972) Tomlinson, M., Woodward, J.: Pile Design and construction Practice, 6th edn. Longman Scientific & Technical, Singapore (2014) Wrana, B.: Pile load capacity – calculation methods. Stud. Geotech. Mech. 37(4), 83–93 (2015) Vardanega, P.J., Pennington, S.P., Kolody, E., Morrison, P.R.J., Simpson, B.: Bored pile design in stiff clay I: codes of practice. Proc. Inst. Civil Eng.-Geotech. Eng. 165(4), 213–232 (2012)

Development of Earth-Based Mortars for Usage in Earth Construction Fady Halim1, Rana Khali1, Safwan Khedr1, Mohamed Darwish2,3(&), Amani Saleh1, Dalia El-Arabi1, Sara Henry1, Ali El-Menoufy1, and Joe Said1 1

American University in Cairo, Cairo, Egypt Ahram Canadian University, Giza, Egypt [email protected] Adjunct Faculty, American University in Cairo, Cairo, Egypt 2

3

Abstract. Earth as a construction material has been used for thousands of years by civilizations all over the world. Earth construction offers a low energy alternative to conventional building materials in addition to being cost effective and safe, qualities which are particularly important with the ever-growing need for increased awareness to reduce energy consumption worldwide. This study is a part of a comprehensive research at the American University in Cairo on earth construction material. The goal of this paper is to test and develop a mortar to bind Compressed Earth Blocks (CEBs) to be used for low-cost housing using materials and techniques which are feasible within rural areas in various parts of the world. This paper includes an experimental program to test 20 different mortar mixes all containing loam and sand testing the effect of adding lime, lignin sulfonate and palm fibers on compressive strength and durability. The variations in strength and durability of the mortars with the different percentages of the mixture components were studied. Based on the results, the best mortar based on compressive strength and durability was selected to be used in construction at a future stage of the experimental program.

1 Introduction The series of scientific discoveries and innovations in the field of construction cause changes in public economic systems for several developed countries. The demand for sustainable construction materials at low cost is growing as social, economic, and environmental issues evolve in societies today. As architectural heritage, earth block masonry attracts the interest of engineers for maintenance and modern construction since it is a material of high environmental and economical profile. Over the past five decades, earthen buildings were the most sustainable and widely used construction materials in developing countries. While our ancestors lived in buildings made of raw materials free from toxics, nowadays, residential buildings contain a high amount of chemicals and heavy metals, that contaminate indoor air, thus causing several health related problems (Chikara et al. 2009) Earth construction is not associated with the adverse effects of indoor air volatile organic compounds (VOCs) so the occupants of these buildings have a superior indoor air quality (Wargocki et al. 2000). © Springer Nature Switzerland AG 2020 P. Pinto et al. (Eds.): GeoMEast 2019, SUCI, pp. 7–17, 2020. https://doi.org/10.1007/978-3-030-34190-9_2

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As there are numerous studies on soil selection, two studies were considered to be the most complete, Delgado et al. (2007) compiled the existing standards on soil selection for compressed earth block (CEB) production and Burroughs (2008) tested over 100 different soils for rammed earth construction (soil for rammed earth is very similar to soil for CEB). Summarizing the findings in these two investigations, soil suitable for CEB production has the following composition: 0–30% of gravel; 25–70% of sand; 20–45% of silt and clay. The soils liquid limit should range from 25 to 50 and its plastic limit from 2 to 30. Water added to the dry mix in CEB production usually ranged from 10 to 13% (Bahar et al. 2004; Rigassi 1985; Riza et al. 2010), normally determined with the Proctor test for optimum moisture content. According to Agha (2003) and Khedr et al. (2003) research mixes revealed optimum water content value of 11%. The water content values plotted against the unconfined compressive strength of the compressed cylinders revealed optimum water content values which can be with an average value of 11% through the range of molding pressure used. Agha (2003) and Khedr et al. (2003) also studied the effect of the percentage of palm fibers added by weight on the results of dry flexural strength tests of CEB. The strength values increased from 0.64 to 1.18 MPa when the quantity of fibers was increased to 0.5% by weight. The blocks keep gaining strength till an optimum value of the fiber is reached, which lies between 0.5% and 1%. The strength value starts to drop to reach 1.21 MPa with a 1% of fibers. The values keep decreasing to reach 1.01 MPa with a 1.5% fiber reinforcement. The curing duration had no effect on the values of the dry flexural strength. Meanwhile the same research showed that high initial rate of strength increase with the increase in molding pressure. At molding pressure of 16 MPa, the gain of strength with increased molding pressure is not significant. Samples molded at pressures less than 4 MPa appeared in a loose texture. This suggested a practical range of molding pressure between 4 to 16 MPa (Agha 2003; Khedr et al. 2003). As it was clear that the blocks compressed using molding pressure below 4 MPa appear in a loose texture lime stabilization was apparently needed. Lime stabilization with 5% by weight was found enough to gain sufficient strength for the blocks manufactured using a high molding pressure to reduce the volume of air voids so the lime and soil will get a closer contact and the stabilizing reactions can take place more easily given that a curing duration of at least 28 days is allowed. The blocks gain its strength through four basic reactions: cation exchange, flocculation and agglomeration, carbonation, and pozzolanic reactions. The long term pozzolanic reaction between the lime and soil particles is believed to be the most important reaction as it occurs between lime and the clay minerals to create a variety of cementitious compounds which bind the soil particles together. Lime also reduce the water absorption for the clay so it enhances the workability with a less water content and make it less sensitive to environmental changes which happen because of the changes in moisture content (Khedr et al. 2003). For blocks manufactured with lower values of molding pressures, it could be estimated that higher values of lime stabilization and curing duration may be needed to achieve similar strength values, yet this point needs further study (Agha 2003).

Development of Earth-Based Mortars

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However, none of the aforementioned researchers worked on the development of mortar made purely from earth materials to be used with CEBs in construction of structures. The main objective of this current research is to fill this gap and develop and test earth-based mortar to be used with compressed earth blocks. Another main paramount of this work was to develop this mortar while keeping the cost as low as possible by means of using local materials containing only natural eco-friendly components. The material properties of twenty different mortar mixes were studied to find out which mix would be the most compatible with compressed earth blocks. The main body of this paper focuses on testing 20 different mortar mixes to be used in earth construction of wall bearing low-rise buildings to attain the same range of strength of CEBs developed by Agha (2003) and Khedr et al. (2003) during a research at the American University in Cairo on earth construction material. The following parameters were acquired from agha’s research and used throughout the study: 11% water content was kept constant within all 20 mortar mixes and a 1:1 ratio of sand-toloam was kept constant for all studied mixtures.

2 Methodology 2.1

Material Procurement

To develop the different mortar mixes, soil or loam similar to that used in the previous research of Agha (2003) had to be acquired. Three soil samples were obtained from different depths from a quarry located in Al-Fayoum, Egypt and tested to choose the one that is similar to that used by Agha (2003). Figures 1 and 2 present the soil gradation curves of soil 1 and soil 3. However, soil 2 was discarded from the tests as it contained a high percentage of sand. This might be linked to the fact that this sample was from a higher altitude than the other samples. According to the results in Table 1, soil 1 is closer in properties to the previous research. Soil 3 had very high plasticity index which indicates that it is organic. So, soil 1 was selected for use in the mortar as it was the nearest to the one used by Agha (2003) and Khedr et al. (2003). 2.2

Sample Preparation

The mortar was divided into four groups with five mixes per group, as shown in Table 2. The ratio of sand to loam was 1:1 and was kept constant throughout all the mixes. Within the four groups, the lime percentages were varied from 0 to 16%. The additives differentiate among the groups. Group A was the control group with no lignin sulfonate and no palm fibers added. Group B contained different percentages of lignin sulfonate additive and no palm fibers. Group C contained different percentages of palm fibers with no lignin sulfonate. Finally, group D contained different percentages of both lignin sulfonate and palm fibers. The lignin sulfonate percentages were determined from initial pilot tests performed. Nine cubes were done for each mix: three for 7 day tests, three for 28 day tests, and three for durability tests. Firstly, the loam and the palm fibers were grinded using the machine shown in Fig. 3. The sand and loam were then sieved on sieve #4. In addition, it was made sure

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100.0 90.0 80.0

% passing

70.0 60.0 50.0 40.0 30.0 20.0 10.0 0.0 10

1

0.1

0.01

0.001

0.0001

0.001

0.0001

Size (mm)

Fig. 1. The gradation curve for soil 1

100 90 80

% Passing

70 60 50 40 30 20 10 0 10

1

0.1

Size (mm)

0.01

Fig. 2. The gradation curve for soil 3 Table 1. Properties of acquired soil samples Soil used by Agha (2003) Soil 1 Soil 3 Specific gravity 2.56 2.35 2.77 Plasticity index 27.2% 31.5% 84.6%

Development of Earth-Based Mortars

11

Table 2. Proportions of the different mortar mixtures studied. Group # Mix # Constituents Loam Sand A A1 50% 50% A2 48% 48% A3 46% 46% A4 44% 44% A5 42% 42% B B1 49% 49% B2 45.5% 45.5% B3 43% 43% B4 40.5% 40.5% B5 38% 38% C C1 49.5% 49.5% C2 47% 47% C3 44.5% 44.5% C4 42% 42% C5 39.5% 39.5% D D1 47.5% 47.5% D2 44.5% 44.5% D3 41.5% 41.5% D4 38.5% 38.5% D5 35.5% 35.5%

Lime 0 4% 8% 12% 16% 0 4% 8% 12% 16% 0 4% 8% 12% 16% 0 4% 8% 12% 16%

Lignin sulfonate Palm fibers 0 0 0 0 0 0 0 0 0 0 4% 0 5% 0 6% 0 7% 0 8% 0 0 1% 0 2% 0 3% 0 4% 0 5% 4% 1% 5% 2% 6% 3% 7% 4% 8% 5%

that the palm fibers do not exceed 15 mm to accommodate to the size of the mortar cubes, Then, a mechanical mixer was used to get an even distribution for all the mortars. The mixes were left in a plastic bag for twenty-four hours to make sure that all the water is absorbed and evenly distributed. After that, the blocks were compacted. The compaction was done using a custom-made steel plate and a steel rod used to compact mortar in 50 mm  50 mm  50 mm molds as shown in Fig. 3. Finally, the cubes were covered in plastic and left for curing for seven days then left to dry for half a day in the sun before testing. 2.3

Compressive Strength Testing

The compressive test was done on all twenty mortar mixes using a compressive strength testing machine. The pressure was applied until the cubes fail and the results were recorded. This was done after 7 and 28 days of curing.

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Fig. 3. The custom-made steel plate and a steel rod used to compact mortar in 50 mm  50 mm  50 mm molds

2.4

Durability Testing

Seven cycles of wetting and drying were performed on the twenty mixes to test the durability. The cubes were placed in the oven for twelve hours then placed in the water for another twelve hours and weighed.

3 Results 3.1

Compressive Strength Tests

Table 3 shows the average 7 and 28-day compressive strength tests. Most of the mixes showed visual trends with a couple of outliers. According to the results, Groups A and B had higher compressive strengths while groups C and D had the lower compressive strengths. This could be attributed to the addition of palm fibers to the mixtures within groups C and D causing a reduction in strength due to the possible volumetric changes caused by these fibers absorption of water within the mixture. The failure pattern exhibited in most of the cubes was a typical brittle failure as shown in Fig. 4 which is in agreement with the failure patterns experienced by most of earth-based materials experimented by the aforementioned researchers. 3.2

Durability Tests

Initially, the volumetric changes that occurred in the durability tests were supposed to be measured. However, most of the cubes disintegrated completely in the water. Therefore, visual inspection became a more reasonable alternative. Groups A and C disintegrated and completely lost their shape after the first cycle, as shown in Fig. 5. On the other hand, groups B and D remained intact throughout the seven cycles, as shown in Fig. 6. This can be linked to the fact that groups B and D contained lignin sulfonate unlike the other two groups.

Development of Earth-Based Mortars Table 3. Results of the compressive strength testing Group 7 - day test (MPa) Group A A1 3.60 3.56 3.68 A2 2.52 2.44 3.16 A3 3.40 3.24 3.08 A4 3.12 3.36 2.68 A5 2.52 3.08 3.12 Group B B1 1.76 1.44 1.44 B2 1.88 1.92 1.80 B3 3.44 3.28 3.28 B4 2.12 2.16 1.60 B5 0.92 1.16 0.96 Group C C1 2.96 3.16 3.08 C2 2.92 3.00 3.20 C3 3.40 3.28 3.16 C4 2.16 3.00 1.72 C5 1.36 1.6 1.28 Group D D1 0.88 0.84 0.88 D2 1.24 1.20 1.12 D3 1.16 1.00 0.96 D4 1.40 1.08 1.48 D5 1.48 1.56 1.40

Average 28 - day test (MPa)

Average

3.60 2.70 3.20 3.10 2.90

A1 A2 A3 A4 A5

3.16 3.48 2.04 1.20 1.56

3.72 3.40 2.40 1.08 1.52

3.32 2.84 2.12 1.28 1.52

3.40 3.20 2.20 1.20 1.50

1.50 1.90 3.30 2.00 1.00

B1 B2 B3 B4 B5

1.68 2.12 3.54 3.16 1.56

1.08 1.76 3.38 3.24 2.16

1.64 2.04 3.38 2.96 1.68

1.50 2.00 3.40 3.10 1.80

3.10 3.00 3.30 2.30 1.40

C1 C2 C3 C4 C5

1.84 1.04 3.22 2.80 1.96

1.36 1.28 3.24 3.00 1.80

1.36 1.56 3.68 2.44 1.76

1.50 1.30 3.40 2.70 1.80

0.90 1.20 1.00 1.30 1.50

D1 D2 D3 D4 D5

1.28 1.24 1.24 1.60 1.40

1.72 1.32 1.72 1.60 1.56

1.12 1.56 1.12 1.24 1.84

1.40 1.40 1.40 1.50 1.60

Fig. 4. The failure pattern exhibited in mortar cubes subject to compression.

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Fig. 5. The shape of groups A and C after the first cycle of wetting and drying.

Fig. 6. Groups B and D after seven cycles of wetting and drying.

3.3

Choice of Mortar

Based on the results of the compressive strengths tests and the durability tests, the best mortar mix out of the twenty mixes was chosen. Since only groups B and D passed the durability tests, groups A and C were discarded. In addition, group B had higher compressive strength than group D, so group D was also discarded. Finally, the optimum mortar was selected using the trend graph shown in Fig. 7. The graph shows that mix B3 has the highest strength corresponding to a moderate percentage of lime of 8%, a moderate percentage of lignin sulfonate of 6% and 43% sand and 43% loam.

Development of Earth-Based Mortars

15

3.5

Compressive Strength (MPa)

3 2.5 2 1.5 1 0.5 0 0

1

2

3 Mix #

4

5

6

Fig. 7. The compressive strengths of group B after 7 days

4 Conclusions The conclusions could be summed up by all the analyses stated above in the following bullet points: • Mortar mixes containing lignin sulfonate have better durability and adhesion but significantly varying compressive strengths depending on the percentage of lignin sulfonate. • Lime tends to coagulate during the curing stage which causes volumetric change during the durability tests, which weakens the mortars’ strength and reduce its durability. This problem was solved with the addition of lignin sulfonate. • Similarities in conduction method, curing, drying, and placing is necessary when creating any object with the blocks for homogeneity in load bearing. • The best mix of mortar is that of group B3 consisting of (by weight): – 43% Loam – 43% Sand – 8% Lime – 6% Lignin Sulfonate

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5 Recommendations 5.1

Future Research

From the results above, four different research recommendations are supplied below: 1. Lignin Sulfonate • Testing the blocks, as well as the mortar, with lignin sulfonate. It is assumed that it will help the mortar durability as well as increase cohesion between the blocks if used to build subsequently in mass production. 2. Sand and Loam Ratio • During the research phase after starting work, it was noticed that some papers suggested a different sand to loam ratio in different areas of the world. Testing this could yield higher strength results and better adhesion of the materials. 3. Drying time • Increasing the drying time of the mortars yielded higher strength so therefore it should be increased in further research. 4. Water in Mortar • An increase in water in the mortar mix could be tested in order to reach the ideal ratio that allows for the highest strength, but also better workability since it was difficult to properly compact the mortar. 5.2

Future Applications

Future applications for compressed earth blocks are endless, since it is already being used with the use of cementitious mortar. The ability to build these blocks from local materials allows for the use of these blocks in rural areas and in the future the ease of construction could allow low income families to build their own houses easily. Also, the industry of compressed earth blocks can be changed from now on to be even more sustainable if the mortar used could protect the environment as well. To reach these future applications a more durable mortar should be tested in order to protect the blocks from any harsh environment it may face. The mortar being properly workable and adhesive to the sides of the mortars is critical to the success of the blocks, as they should cover the blocks completely and be used as a substitute for plastering the blocks. Acknowledgments. We would like to thank the lab personnel at the construction engineering department at the American University in Cairo for their constant help and support throughout the process of this work. Also we would like to thank Dr. Naahed Zarifah and Ms. Dalia Zarifah for their help in providing the lignin sulfonate used within the study.

Development of Earth-Based Mortars

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References Bahar, R., et al.: Performance of compacted cement-stabilized soil. Cem. Concr. Compos. 26, 811–820 (2004). https://doi.org/10.1016/j.cemconcomp.2004.01.003 Burroughs, S.: Soil property criteria for rammed earth stabilization. J. Mater. Civ. Eng. 20(3), 264–273 (2008). https://doi.org/10.1061/(ASCE)0899-1561(2008)20:3(264) Delgado, M.C., et al.: The selection of soils for un-stabilized earth building: a normative review. Constr. Build. Mater. 21, 237–251 (2007) Khedr, S., Abou-Zeid, M.N., Agha, S.: Concerns in manufacturing compressed earth blocks. In: Proceedings of the 35th Canadian Society of Civil Engineering Annual General Conference, Moncton. CSCE, Canada (2003) Pacheco-Torgal, F., Jalali, S.: Toxicity of building materials: a key issue in sustainable construction. Int. J. Sustain. Eng. 4(3), 281–287 (2011). https://doi.org/10.1080/19397038. 2011.569583 Rigassi, V.: Compressed Earth Blocks : Manual of Production. CRAterre-EAG, GATE, vol. I (1985) Riza, F.V., et al.: A brief review of compressed stabilized earth brick (CSEB). In: CSSR 2010 International Conference on Science and Social Research (CSSR), pp. 999–1004 (2010). https://doi.org/10.1109/cssr.2010.5773936 Wargocki, P., Wyon, D., Sundell, J., Clausen, G., Fanger, P.: The effects of outdoor air supply rate in an office on perceived air quality, sick building syndrome (SBS) symptoms and productivity. Indoor Air 10(4), 222–236 (2000). https://doi.org/10.1034/j.1600-0668.2000. 010004222

Experimental Investigations of Capacity Response of Root Piles on Combination of O-Cell Test and Conventional Head-Down Test Methods Xiao-juan Li1,2, Guo-liang Dai1(&), Wei-ming Gong1, and Ming-xing Zhu1 1

2

Geotechnical Engineering, School of Civil Engineering, Southeast University, Nanjing, Jiangsu, China [email protected],[email protected], [email protected] Geotechnical Engineering, Lyes School of Civil Engineering, Purdue University, West Lafayette, IN, USA [email protected]

Abstract. This paper introduces installation procedures of root piles, and then presents the results of an experimental study on the behavior of a root pile in Chizhou Yangtse River Bridge. In order to investigate the capacity response of root piles and the effect of roots on shaft resistance, O-cell test and conventional head-down test were conducted on a root pile, with diameter of 1.8 m and length of 49 m. The pile was instrumented with strain gauges along the steel bars of pile in load tests to measure the load distribution along the length of the test pile, as well as its shaft resistance. The research demonstrates that the existence of roots affects the value of average compressive stiffness in O-Cell test, and also changes the transform character of axial loads significantly; the difference between the shaft resistance with roots and that without roots is increase with the increase of load; the existence of roots increased shaft resistance at pile section with roots, but it decreased shaft resistance at the adjacent segments without roots, These positive and negative effects are influenced by the distribution of roots and loading strategy.

1 Introduction Root foundation has been successfully used in some projects, such as Maanshan Yangtse River Bridge, Huainan bridge, Wangdong Yangtse River Bridge, Chizhou Yangtse River Bridge in China for ten years. Unlike ordinary abnormity deep foundation (tapered piles, squeezed branch piles, helical piles or belled piles, etc.), it has many roots around the pile shaft, which is similar as the roots of a plant (Figs. 1 and 2). The load is transferred from the pile shaft to the surrounding soils through these roots. Some of root foundation have large diameters as much as 6 m and hollow inside, which is named as “root caisson” and can be designed by the principle of conventional caisson. While others have smaller section and much slender body which should be considered as deep pile foundation, will be studied in this paper. © Springer Nature Switzerland AG 2020 P. Pinto et al. (Eds.): GeoMEast 2019, SUCI, pp. 18–39, 2020. https://doi.org/10.1007/978-3-030-34190-9_3

Experimental Investigations of Capacity Response of Root Piles

19

Fig. 1. Comparison of tree root to the root pile

Fig. 2. Roots around the pile shaft.

Understanding the response of root foundations subjected to axial load is essential for the development of design methods. Behavior of root caisson have been explored by many field tests or finite-element methods (Yin et al. 2009; Mu et al. 2010; Hu 2010; Bai 2013; Gong 2008; Zhang 2009; Gong 2015), results show that roots can increase vertical bearing capacity of caisson significantly. Yet, to our best knowledge, no results of static load test offering reliable capacity response on root piles are available in the literature. However, the uncertainty of this positive effect on piles leads

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to over-conservation design in real projects. Therefore, it is important to reveal the effect of roots on shaft resistance in root piles. Both the conventional top-down static load test (kentledge test) and O-Cell test are reliable approaches (Russo 2013) to study the response of piles under axial loading. However, static loading test is not substituting for determining the behavior of root piles for that the load test reaction systems often do not have the capacity to take root piles all the way to plunging. Furthermore, in O-Cell test, several hypotheses are required to transform experimental results into those of an equivalent conventional load test. But the transform procedure is not clear in root piles yet, even it has been successfully applied in some root caissons with very large bearing capacity (Yin et al. 2009; Gong 2008; Zhang 2009; Dai 2012; Gong 2015). Therefore, special combined load tests—the bi-directional test and the top-down load test–was used in a root pile in Chizhou Bridge project. Unlike the method introduced by Fellenius et al. (Fellenius 2010), there are two procedures in this research generally: (1) in the O-Cell test, a load cell was used in pile body near the pile end to push two parts of the pile to opposite directions, until the ultimate shaft resistance of pile above load cell reached; (2) after three months, the top-down load test was applied to push the upper part of the pile back. Therefore, the ultimate shaft resistance of the pile above the load cell can be obtained, as well as the performance of roots. This paper reports on the instrumentation, installation, and field tests of a root pile in Chizhou Bridge. The pile with a diameter of 1.8 m had a depth of 49 m in a multilayered soil profile. Load-settlement, load-transfer curves and unite shaft resistance were obtained from the related load tests. In addition, the shaft resistance with root keys was compared with its counterpart that without roots. Finally, the enhancement of roots as well as their mechanics characters were discussed.

2 Test Site: Subsoil Conditions and Layout of the Tests 2.1

Site Description

A cast-in-situ root pile was load tested by the O-Cell test and kentledge test in Chizhou Yangtse River Bridge, China. The test pile will be used in the foundations of a twospan bridge. The deposition of sediments there are influenced by Yangtse River, so the soil profile where the pile located was mainly silty clay and silty sand with relative uniform size and high roundness. A boring hole were performed close to the test pile location before the O-Cell test was applied (Fig. 3). Tables 1 and 2 give the detail information of soil profile at the test site, Fig. 3 shows the results of SPT, N refers to the average SPT blow counts in 30 cm at every depth. The soil profile consists primarily of layers of silty clay, medium dense to dense fine sand, gravel mixtures and moderate weathering conglomerate. The water table was found at 2.7 m below the ground. The relative density of the sand layer (DR) is not given directly. According to the method proposed by Meyerhof (1957) and Skempton (1986), the relation density in terms of N is given as:

Experimental Investigations of Capacity Response of Root Piles

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Fig. 3. Schematic profile of soil conditions and test piles Table 1. Parameters of clay Depth (m) 0–3.0 3.0–5.1 5.1–6.8 6.8–9.0

Soil Silty clay Silty clay Silty clay Silty clay

Density q (g/cm3) 1.81

su, NC (kPa) 2.72 0.972 34.0 38.8 23.0 0.75 15.8 13.6

1.77

2.72 1.202 43.3 41.4 22.2 1.10 19.2 17.7

1.80

2.72 0.943 28.0 27.7 17.1 1.03 10.6 25.5

1.85

2.72 0.978 34.5 35.5 21.6 0.93 13.9 31.2

Gs

e

w

LL

vffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi u DR N u ¼t r0 100% A þ BC v

PL

LI

PI

ð1Þ

pA

where A ranges from 27 to 46 (36.5 was used in this paper), B is approximately 27; 0 C = 1 for NC soils; rv is the vertical effective stress of soil; pA = 100 kPa. The result of DR is shown in Fig. 3.

22

X. Li et al. Table 2. Parameters of sand and gravel of particle composition (%) uc (°) under 2– 0.5– 0.25– 20 20– 2 0 0

Silty sand

1.90

0

0

0

8.5

76.1

15.4

27.6

Silty sand

1.91

0

0

0

9.4

76.3

14.3

28.0

Silty sand

1.93

0

0

0

8.6

76.2

15.2

28.0–30.1

Round gravel mixtures

1.98

0

57.5 20.1 9.4

4.2

8.8

2.2

–

Test Pile Instrumentation

The test pile is a cast-in-situ concrete pile with a diameter of 1.8 m and embedment depth of 49 m (Fig. 3). The main shaft of test pile has a cross-sectional area of 2.5434 m2 (Ac) with concrete elastic modulus of 300 MN/m3 (Ec) and concrete density of 2550 kg/m3. The main shaft of pile was installed precast roots at different depth (Fig. 4), with 4 roots in each layer, and 24 layers in total. The layout of roots in every two adjacent layers is shown in Fig. 4(b) to prevent interferences of soil resistance among root layers. Every root has a cross-sectional size of 0.15  0.15 m2 and was stretched into soil of 0.45 m. The roots were well reinforced to avoid crushing or yielding. An instrumentation scheme with vibrating-wire strain gauges was used to measure the axial deformation of the test pile and to obtain the load transfer curves. The dynamic range of the strain gages was 0–100 kN of tension and 0–50 kN of compression, respectively. Strain gauges were attached at four orientations at 37 different depths along the pile length. The layout of the strain gauges is shown in Fig. 4(a), (c) and (d). However, some of them were cut off or work in poor condition after pile construction. Fortunately, at least one or two strain gauges show good performance at each depth. A 0.71 m-high load cell was used as loading device in O-Cell test, it was wielded as an assembly with the reinforcement cage (Fig. 5(a)). The load cell located at 5 m above the pile base (Fig. 3). The diameter of the load cell is the same as the reinforcement cage. The displacement signal of upper plate and lower plate of cell was monitored by displacement sensors and recorded by data acquisition device. The construction of the root pile generally contains four procedures (see Fig. 5(b)): (1) drill a hole using the slurry method, (2) put down a rebar cage, (3) install root keys at each depth, (4) fill the hole with concrete. As shown in Fig. 5(c), at the stage of rebar cage manufacture, preserved guide holes are on the cage for the installation of root keys.

Experimental Investigations of Capacity Response of Root Piles

23

Fig. 4. (a) The schematic diagram of pile, the layout of (b) roots and (c)-(d) strain gauges on rebar cage.

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Fig. 4. (continued)

At the stage of roots installation, an installation machine pushes the roots through guide holes into soil to required length. This procedure is valuable, which can ensure the roots in predefined location. The device is suitable for construction of root piles with diameters of 1.5 m, 1.8 m and 2.5 m. 2.3

Load Testing Method

The O-CELL tests and the top-down load test were conducted on the same pile. The purpose of load tests is to confirm the behavior of pile and root keys. The test program included three stages. Stage 1: As shown in Fig. 6(a), activated O-Cell in increments of 1000 kN applied every 120 min. The top and bottom planes of load cell were apart and pushed their corresponding pile sections to move contrarily until the ultimate resistance of the pile above the load cell reached, the expansion of load cell also reach to an impressive

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25

Fig. 5. (a) Rebar cage with guide holes, (b) the installation procedure of root piles and (c) guide holes.

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Fig. 6. Load tests procedures for SZ1 and SZ2 (a) O-Cell test, (b) begain of the top-down load test,(c) pushing the upper pile back.

extend and then unloaded the cell. The displacement VS applied load of top and bottom planes were measured at each loading step. Stage 2: Rested the test pile for 3 mouths with the load cell open and free, so that the residual loads in the pile can be fully demobilized and the soil disturbance from OCell test can be restored. Stage 3: loaded pile head in increments of 1000 kN applied every 120 min until pushing the pile above the load cell back. As illustrated in Fig. 6(b) and in Fig. 6(c), the pile above the load cell performed as a shaft-bearing pile free of toe resistance. so that all root keys around pile shaft above the load cell can be fully mobilized in topdown load test. Then unloaded the pile.

3 Test Site: Subsoil Conditions and Layout of the Tests 3.1

Load-Displacement Response

In most load tests, the 0.1B settlement criterion is used to define an ultimate limit stage of pile, which depends on the ultimate unit base resistances, Therefore, the displacement criterion is reasonable for base resistance of roots. The displacement response of O-CELL test at stage 1 is shown in Fig. 7. The solid line at the positive coordinate axis is the load-settlement curve of top plane of load cell

Experimental Investigations of Capacity Response of Root Piles

27

(Qupper-supper), and the dash line is that at the pile head (Qupper-stop), the negative displacement is from the bottom plane of load cell (sdown). From Fig. 7, the loadsettlement response changed smoothly in the beginning of the loading up to an applied load of 1,3000 kN (corresponding to supper of 20.45 mm and stop of 18.66 mm), after which the slope of the load-settlement curve decreased significantly. The ultimate load, corresponding to a supper of 47.51 mm and stop of 45.61 mm, was 1,4000 kN. Finally, supper of 39.71 mm and stop of 37.89 mm remained after unloading. On the other hand, for the downward one (sdown), it reaches to 14.01 mm at load of 1,4000 kN slowly and decreases to 9.82 mm when unloaded.

Fig. 7. The load variation of load-cell and pile-head displacement of SZ2 (stage 1).

The displacement response of test pile in stage 3 are shown in Fig. 8. The loadsettlement response was almost linear in the beginning of the loading up to an applied load of 1,5000 kN (corresponding to settlement of 52.8 mm). after that, an extra load of 200 kN was applied, but no incremental settlement existed at the pile head. Finally, settlement of 42.73 mm remained after unloading. There are two reason to confirm that the upper part of test pile are likely contacts with the lower part: (1) there is no incremental displacement at load of 1,5200 kN, which does not accord with the incremental tendency of settlement from load of 2,000 kN up to 1,5000 kN; (2) the residual distance between the upper part and lower part of pile is 49.53 mm at unloading stage in Fig. 7, and residual settlement of pile head is 42.73 mm at unloading stage in Fig. 8, there is 6.8 mm difference, which can be caused by the strain recovery of soil or the dead-load of pile.

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Fig. 8. The load variation of pile-head displacement of the test pile (stage 3).

3.2

Compression Stiffness

Compression stiffness is a very important factor that influence the load-transfer curves of pile. However, cross section in the root pile is not uniform, the areas are different at different pile length segments (Fig. 4). Therefore, the average compression stiffness of root piles was discussed in this section. Figure 7 present the displacement of the load cell and the pile head at different load step in field tests, respectively, the movement of pile head is also measured, thus, the pile compression Δsi can be calculated by the subtraction of supper and stop at load step i. Δsi can be calculated from: Dsi ¼ supper;i  stop;i

ð2Þ

where i is the different loading level. As for the force condition of the pile in O-CELL test, the shaft resistance distribution is assumed to be trapezoid along with the depth of soil layer. The average compression stiffness of pile segment above load cell can be determined by plugging Eq. (2) into Eq. (3): Ep Ap ¼

2ðQi  WÞLupper 3ðsupper;i  stop;i Þ

ð3Þ

where Lupper is the length of upper pile (m), Qi is the load at i loading level (kN) and EpAp is the compression stiffness of upper pile (kN).

Experimental Investigations of Capacity Response of Root Piles

29

Though the existence of root keys along and around the pile shaft makes rook pile difference from traditional pile at some extend, the overall tendency of shaft resistance of soil is similar along the pile. Thus Eq. (3) can be used in root pile foundation. To study the compressive stiffness of root piles, A complete non-dimensional framework is used, as a is the value of average compression stiffness divided by EcAc, and Qi is divide by W. where EpAp is the result from Eq. (3), the elastic modulus of reinforced concrete of pile which recommended value in code multiplying its section area, here EcAc = 7.63 * 107 kN, W = 1679 kN. Another root pile near the test pile was also tested by same load test methods in the same site with the dimeter of 1.6 m, length of 49 m and 22 layers of roots. Because most its strain gauges performed not well during load tests, it cannot obtain the detail performance of axial force. However, the values of supper, sdown and stop in O-Cell test have been recorded with good accuracy, therefore, these values can be also used to discuss compression stiffness of root piles. The value of a of two root piles is plotted in Fig. 10, where the abscissa is applied load divided by dead load of the pile segment above load cell. Three points has been removed from Fig. 10 because of significant deviation of overall trend. From least square fitting method, a has positive relationship with applied load, the fitting curve is a = 2(Qi/W)1.089, the fitting error is 0.84, and the average value of a is 1.118. This means that, the average compressive stiffness of root piles in O-Cell test is related to the applied force, which is not a commonly believed constant. In addition, the average stiffness is larger than that of concrete stiffness of the truck of pile. The main reason is the existence of roots around pile shaft. qs is the unit shaft resistance of pile segment between strain gauges layers. In Fig. 9(a), there is no roots in pile segment, qs is caused by shear friction along pile shaft, while in Fig. 9(b), there has roots, so qs is the sum of shear friction of the pile segment and base resistance of roots. The existence of roots can enlarge the contact area of soil and provide more soil resistance. With the load applied increased, the incremental soil resistance caused by roots increased as well, this would result increase of compressive stiffness.

Fig. 9. Unit friction resistance of pile section (a) without root keys and (b) with root keys

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Fig. 10. The force condition of upper pile.

3.3

Shaft Capacity

Vibrating-wire strain gauges mentioned above in this paper recorded the strain changes at 37 different locations along the test pile length during O-Cell test and kentledge test, these data were used to obtain the profiles of the axial load distribution along the pile segment above load cell at each loading level. Figure 11 shows the depth profiles of the axial loads (load-transfer curves) obtained from the strain gauges along pile at each loading level. From Fig. 11(a), at the load level of 2000 kN, axial load existed in a small area near the load cell, the transform of axial force is mainly obstructed by the last three layers of roots. The load transferred to the pile head at the applied load of 4000 kN up to 1,4000 kN. From Fig. 11(b), the axial load transferred to the load cell at the applied load of 7000 kN up to 1,5000 kN. From Fig. 11, the existence of roots resisted the transform of axial loads significantly. When the vertical load was applied, the small load was fully absorbed by shaft resistance along a section of pile near load applied device. root keys near load device worked at first, then the farther layers worked, successively. The point which axial force is zero indicates no load is transferred from the pile shaft to the soil. As the load is increased, the point moves down (or upward in O-Cell test) until shaft resistance is fully mobilized. The average unit shaft resistance between two neighboring strain gauges can be calculated by the difference between the axial loads divided by the area of the outer surface of the pile section between the two gauges. Thus, the slope of these curves in Fig. 11 indicates the values of the shaft resistance at different load levels. The profile of the calculated unit shaft resistance in O-Cell test and in kentledge test at different load levels versus depth is plotted in Fig. 12, respectively.

depth/m

depth/m

Experimental Investigations of Capacity Response of Root Piles

0

2000kN

-5

3000kN

-10

4000kN

-15

5000kN

-20

6000kN

-25

700kN

-30

8000kN

-35

9000kN

-40

10000k N

-45 0

5000

10000 15000 axial force/kN (b)

20000

Fig. 11. (a) The Axial force of SZ2 in (a) O-CELL test and (b) head-don load test

31

32

X. Li et al.

Fig. 12. Friction resistance of the test pile at different load levels in (a) O-CELL test and (b) the top-down load test

Experimental Investigations of Capacity Response of Root Piles

33

From Fig. 12, the mechanical behavior of pile segments with roots is different from that without roots, the former had larger shaft resistance than the latter. This is because the former is made of two parts: (1) the sum of friction along truck of pile and that along the sides of roots, (2) the compressive resistance at the contact of root base with the underlying soil. While the latter is mainly from the friction along pile shaft. The existence of (2) is the mainly difference from traditional pile without roots. From Fig. 12, we can also find that the difference between the shaft resistance with roots and that without roots was small at small load, and as the load increased, the difference became more obvious. The reason is that friction resistance of pile can be fully mobilized for small pile displacements (4.5 mm–18 mm in this paper), while base resistance of roots needs larger displacements (at least 45 mm in this paper). For pile segments with roots at small load, the friction at pile shaft is mainly resistance, the contribution of base resistance of roots is small, thus, the difference is small between pile segments with roots and that without roots; as the load is increased, the base resistance of roots would be gradually mobilized, which results larger difference between the shaft resistance with roots and that without roots. There are two reason of confirming unit shaft resistance at last load level in Fig. 12 reached to the ultimate value: (1) the settlement of pile segment above load cell is enough for friction resistance of pile-soil interaction be fully mobilized, and (2) it is also enough for the base settlement of roots reach to 10% of their length or width. 3.4

Comparison of Ultimate Shaft Resistance with Different Methods

To find out the difference of shaft resistance between the root pile and a traditional pile without roots, the shaft resistance with roots or without were also compared by different calculation methods. The equations for shaft resistance of nondisplacement piles is as following: In sand (Fleming 1992): 0

qsL ¼ Krv tan d

ð4Þ

In clay (O’Neill and Reese 1999): qsL ¼ asu

ð5Þ

Where qsL is the ultimate shaft resistance; K = 0.7; d is the friction angle of interface, d = uc, uc is determined by quick shear tests; a = 0.55; su is the saturated shear strength. The equations for shaft resistance of nondisplacement piles is given as: In sand (Salgado 2006a; Salgado 2006b): 0

qsL ¼ Krv tan d K ¼ 0:7 exp K0

  0   r 0:0114  0:0022 ln v DR pA

ð6Þ ð7Þ

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X. Li et al.

In clay (Hu and Randolph 2002, Salgado 2006a): qsL ¼ asu h i a ¼ 0:4 1  0:12 lnðpsuA Þ

ð8Þ 3  OCR  5

ð9Þ

Where DR is the relative compactness of sand; r′h is the horizonal effective stress of 0 soil, r′h = K0rv ; pA = 100 kPa; K0 = 0.45; qb,10% is the base resistance at settlement of 10% of base diameter. a is conservative in NC soils. Tables 1 and 2 summarizes the values of the input variables that were used in pile resistance estimations. Ultimate shaft resistance from Eqs. (4) and (5), Eqs. (6) and (8) are presented in Fig. 13.

Fig. 13. Ultimate shaft resistance in different methods

In Fig. 13, the magnitude of the unit ultimate shaft resistance from z = –25.7 m toward z = 0 m by different methods is presented. Equations (4) and (5) gave overestimated results of total soil resistance. Equations (6) and (8) gave much better prescription. Comparing with the measured ultimate shaft resistance at pile segments which roots existed, the unit ultimate shaft resistance which roots did not exist are much smaller in both load tests, even they were at similar depth. What’s more, the former is slightly larger than the result from Eqs. (6) and (8).

Experimental Investigations of Capacity Response of Root Piles

35

Table 3. The effect of roots on ultimate shaft resistance Soil

Depth (m)

Silty clay and silty sand Silty sand Silty sand Silty sand

0.00–11.04 (no roots) −11.04–−11.64 −11.64–−12.54 −12.54–−13.64 (no roots) −13.64–−14.24 −14.24–−15.14 −15.14–−16.24 (no roots) −16.24–−16.84 −16.84–−17.74 −17.74–−19.45 (no roots) −19.45–−20.34 −20.34–−21.44 (no roots) −21.44–−22.34 −22.34–−22.94 −22.94–−24.04 (no roots) −24.04–−25.54 −25.54–−26.64 (no roots) −26.64–−26.84 −26.84–−28.14 −28.14–−29.24 (no roots) −29.24–−29.84 −29.84–−30.67 −30.67–−31.84 (no roots) −31.84–−32.44 −32.44–−33.04 −33.04–−34.44 (no roots) −34.44–−35.04 −35.04–−35.94 −35.94–−36.54

Silty sand Silty sand Silty sand Silty sand Silty sand Silty sand Silty sand Silty sand Silty sand Silty sand Silty sand Silty sand Silty sand Silty sand Silty sand Silty sand Silty sand Silty sand Silty sand Silty sand Silty sand Silty sand Silty sand Silty sand Silty sand

O-Cell test (kN/m2)

Top-down test (kN/m2)

23

25

Percentage (%) OTopCell down test test – –

26 42 30

30 46 36

13 40 –

20 28 –

47 49 34

51 54 43

57 44 –

42 26 –

53 53 32

58 63 47

56 40 –

35 25 –

55 34

65 41

62 –

59 –

54 56 35

67 69 41

59 60 –

63 68 –

65 49

75 58

33 –

29 –

71 71 52

82 83 62

45 37 –

41 34 –

84 84 49

87 90 62

62 71 –

40 45 –

83 83 53

91 93 72

69 57 –

47 29 –

85 85 85

96 97 95

60 – 67

33 – 32 (continued)

36

X. Li et al. Table 3. (continued)

Soil

Depth (m)

Silty sand

−36.54–−37.04 (no roots) −37.04–−37.94 −37.94–−39.64 (no roots) −39.64–−40.54 −40.54–−41.14

Silty sand Silty sand Silty sand Round gravel mixtures

51

72

Percentage (%) TopOdown Cell test test – –

86 56

97 83

69 –

35 –

102 141

131 157

82 –

58 –

O-Cell test (kN/m2)

Top-down test (kN/m2)

From Fig. 13, the exist of roots can increase the shaft resistance of pile, and it is interesting to find that they also have negative effect on the soil resistance of adjacent pile segments without roots, because the soil resistance of pile segments without roots was rather smaller than the results from different methods. This phenomenon existed in both the top-down test and O-Cell test. The reason is that, in a root pile, the calculation unit shaft resistance at a certain depth is the different from that of the traditional cast-in-site piles with uniform crosssectional area, it contains the shaft resistance for the truck of root pile as well as the side of roots, and the base resistance of four roots around pile shaft. On the other hand, with the mobilizing of the base resistance of roots, the pile segments have larger “ability” to bear axial force, this would decrease the value of K of adjacent pile segments without roots, where they had less ability to press surrounding soil. Some pile segments between strain gauges layers have roots, while some do not have. The shaft resistance with roots were compared with that without roots from adjacent segment, so that the effect of roots in a same pile can be studied. The results of ultimate shaft resistance and increase percentage is shown in Table 3, where the ultimate shaft resistances in O-Cell test are the values at load level of 14000 kN, and those in top-down load test are the values at load level of 15000 kN. In Table 3, the positive effects of roots on unit ultimate resistance in O-Cell test were given. Near the pile top, the increase effects of roots were smaller than that at middle part of pile and that near the load-cell. The largest value existed near the load cell. The positive effects of roots on unit ultimate resistance in top-down load test were also given. Near the pile top, the increase effects of roots were smaller than that at middle part of pile, and the effects near the load-cell were also smaller than that at middle part of pile. The largest value existed at middle part of pile. The effects of roots on shaft resistance in O-Cell test were more remarkable than that in top-down load test. This means that different loading strategies influence the distribution of the effect. While near the soil ground, the increase effect is small in both tests.

Experimental Investigations of Capacity Response of Root Piles

37

The increase effects of roots in Table 3 reflected not only the positive effects of roots, but also the negative effect on the adjacent segments without roots. However, it is hard to conclude that the overall effect on total shaft resistance is bad, because we don’t know the real profile of the unit shaft resistance of the pile with same conditions but no roots. While form exists studies of root caisson (Yin 2009; Gong 2008; Zhang 2009; Dai 2012; Gong 2015), where total shaft resistance before roots installation and after installation in same piles were studied, the existence of roots has good effect on the total shaft resistance. 3.5

Contribution of Root Keys

As shown in Eq. (10), the ultimate bearing capacity Qult of a single pile contains two parts: ultimate base resistance Qb,ult and the limit shaft resistance QsL. Qult ¼ Qb;ult þ QsL

ð10Þ

While in root piles, the limit shaft resistance QsL is (1) the sum of shear resistance of truck of pile and side resistance of roots and (2) the ultimate base resistance of roots. Therefore, Eq. (10) can be rewritten as: QsL ¼

n X

ai qsL;i Asi þ

i¼1

n X

bi qb;ult;i Abi

ð11Þ

i¼1

Where qsL,i is the sum of shear resistance of the truck of pile and side resistance of roots with layer i; qb,ult,i is the ultimate base resistance of four roots with layer i. ai and bi are the coefficients of correction, respectively, 0 < ai < 1, 0  bi < 1(if there is no roots with layer i, bi = 0); Asi is the area of the truck of pile and side resistance of roots (if exist); Abi is the area of four root base with layer i. In Eq. (11), ai and bi are there to modify the calculated value of qsL,i and qb,ult,i, respectively. It is difficult to verify the values of ai and bi in this paper by results of the test pile. While their values are influenced by the position where roots exist and loading strategy. In top-down load test, their value would be smaller near ground soil as well as near pile end, and they should be larger at the middle part of pile. In O-Cell test, their value should be smaller near ground soil and be larger at deeper depth. In the O-Cell load test, the value of QsL is different from that of a top-down load test, because their stress state is different. There should be a conversion coefficient c (Osterberg 1989; Mission and Kim 2011; Li 2016a; Li 2016b), which is shown in Eq. (12). The value of c here is 0.82. QsL ¼ cð

n X i¼1

ai qsL;i Asi þ

n X

bi qb;ult;i Abi Þ

ð12Þ

i¼1

Besides influence factors mentioned above, the values of ai and bi are also influenced by embedment length of roots, distance between root layers and soil type. Study results of squeezed branch piles (Qian 2003; Wang 2004; Lu et al. 2004) show that the

38

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increase embedment length of roots increase the total shaft resistance of pile, and narrow distance of root layers have negative effect on total shaft resistance of pile. These will give good reference of determining values of ai and bi.

4 Summary and Conclusions The paper presented results of a top-down load test and an O-Cell test on a cast-in-situ pile with roots in Chizhou Yangtse River Bridge. Because the test pile was fabricated in two segments by load cell device, and the roots were prefabricated, a special procedure was introduced to explain the manufacturing method. The test pile was fully instrumented with vibrating-wire strain gauges to measure the axial deformation along the test pile. Both an O-Cell test and a top-down load test were performed on the test pile. The test pile load-settlement response, load-transfer curves, and limit shaft resistance profiles were obtained from the data measured by the strain gauges. Based on the in-situ testing results presented, the following conclusions can be drawn: (1) The existence of roots affects the value of average compressive stiffness of root piles in O-Cell test. it is not a constant, and it increases with the increase of load level; the average value is 1.118EcAc. (2) The existence of roots resisted the transform of axial loads significantly. Root keys near the load applied device worked at first, then the farther layers worked, successively. (3) In the same pile, the mechanical behavior of pile segments with roots is different from that without roots. The shaft resistance of pile segments with roots had larger shaft resistance than that without roots. (4) The difference between the shaft resistance with roots and that without roots is small at small load, and as the load is increased, it becomes more obvious. (5) The existence of roots not only has the positive effects on shaft resistance at pile section with roots, but also has a negative effect on the adjacent segments without roots. (6) These effects are influenced by the position where roots exist, loading strategy, embedment length of root keys, distance between root layers and soil type. Acknowledgments. This work was supported by China Scholarship Council, National Key Research Program of China (2017YFC0703408) and National Natural Science Foundation of China (51878160) and (51808112), Jiangsu Fundamental Research Program (Natural Fund Project) (BK20180155).

References Bai, R.F.: Numerical Simulation Analysis on Vertical Bearing Mechanism of Root-Caisson. M.S. thesis. Southeast University (2013)

Experimental Investigations of Capacity Response of Root Piles

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Fellenius, B.H., Tan, S.A.: Combination of O-cell test and conventional head-down test. In: Art of Foundation Engineering Practice, pp. 240–259 (2010) Fleming, W.G.K.: A new method for single pile settlement prediction and analysis. Geotechnique 42(3), 411–425 (1992) Dai, G.L.: Application of bi-directional static loading test to deep foundations. J. Rock Mech. Geotechn. Eng. 4(3), 269–275 (2012) Gong, W.M., Hu, F., Tong, X.D., Yin, Y.G.: Experimental study on vertical bearing capacity of root foundation. Chin. J. Geotech. Eng. 30(12), 1789–1795 (2008) Gong, W.M., Wang, L., Yin, Y.G.: Applied and experimental study on the root foundation in the thick covering stratum region. China Civil Eng. J. 48(suppl 2), 69–75 (2015) Hu, F.: Theoretical and experimental research on load bearing characteristic of root-caisson, Ph. D. thesis, Southeast University (2010) Hu, Y., Randolph, M.F.: Bearing capacity of caisson foundations on normally consolidated clay. Soils Found. 42(5), 71–77 (2002) Meyerhof, G.G.: The ultimate bearing capacity of foundations on slopes. In: Proceedings of 4th International Conference on Soil Mechanics and Foundation Engineering, vol. 1, pp. 384–386 (1957) Mission, J.L.C., Kim, H.K.: Design charts for elastic pile shortening in the equivalent top– download–settlement curve from a bidirectional load test. Comput. Geotech. 38(2), 167–177 (2011) Li, X.J., Dai, G.L., Gong, W.M., Xu, W.X., Wang, L.: The research on conversion factor of selfbalanced loading test in sandy area. Rock Soil Mech. 37(sup1), 659–667 (2016a) Li, X.J., Chen, X.J., Dai, G.L., Gong, W.M.: Research of conversion coefficient of cast-in-situ pile at clay area in self-balanced loading test. Rock Soil Mech. 37(sup1), 226–232 (2016b) Lu, C.Y., et al.: Testing study on effect of different soil layers on load transfer of model piles with branch plates. Chin. J. Rock Mech. Eng. 20, 033 (2004) Mu, L.L., Huang, M.S., Gong, W.M., Yin, Y.G.: Response analysis of anchorage foundation under lateral loading. Rock Soil Mech. 31(1), 287–292 (2010) Osterberg, J.: New device for load testing driven piles and drilled shaft separates friction and end bearing. Piling Deep Found. 1, 421–428 (1989) Qin, D.: Engineering application study of squeezed branch pile with high antipulling behavior. Chin. J. Rock Mech. Eng. 4, 045 (2003) Russo, G.: Experimental investigations and analysis on different pile load testing procedures. Acta Geotech. 8, 17–31 (2013) Salgado, R.: The role of analysis in non-displacement pile design. In: Modern Trends in Geomechanics, pp. 521–540. Springer, Heidelberg (2006a) Salgado, R.: Analysis of the axial response of non-displacement piles in sand. In: Geomechanics II: Testing, Modeling, and Simulation, pp. 427–439 (2006b) Skempton, A.W.: Standard penetration test procedures and the effects in sands of overburden pressure, relative density, particle size, ageing and overconsolidation. Geotechnique 36(3), 425–447 (1986) Technical Code for Building Pile Foundations, JGJ94-2008 Wang, D., et al.: Law of load transmission of squeezed branch piles and it’s research advances. Chin. J. Rock Mech. Eng. 23, 4645–4648 (2004) Yin, Y.G., Sun, D.H., Gong, W.M.: Experiment and numerical simulation of the bearing characteristics of root foundations. China Civil Eng. J. 42(12), 162–169 (2009) Thiyyakkandi, S., Michael, M.V., Bloomquist, D., Lai, P.: Experimental study, numerical modeling of and axial prediction approach to base grouted drilled shafts in cohesionless soils. Acta Geotech. 9, 439–454 (2014)

Finite Element Modeling of Soil Arching in Pile Supported Embankment: 2D Approach Naveen Kumar Meena, Sanjay Nimbalkar(&), and Behzad Fatahi School of Civil and Environmental Engineering, University of Technology Sydney, Broadway, Ultimo, NSW 2007, Australia [email protected], {Sanjay. Nimbalkar,Behzad.Fatahi}@uts.edu.au

Abstract. The presence of soft soil located along the coastal belt of Australia poses a serious challenge to the construction of infrastructure projects. The pilesupported embankments provide a viable alternative to deal with such problematic soil. The soil arching mechanism plays a key role in transferring loads to piles in an efficient manner. In this study, a simplified two-dimensional finite element approach is used to demonstrate soil-arching mechanism in a granular embankment supported by piles. The influence of the piled-embankment properties such as elastic modulus of pile, friction angle and modulus of embankment-fill are investigated on the soil arching in term of stress concentration ratio. It is found that the pile and embankment modulli and friction angle significantly affects the arching mechanism. The thickness of arching zone is found to maximum at the mid of subsoil whereas minimum on the pile head and is analogous to the multi-shell theory. The effect of embankment height and pile spacing have also been reported. Keywords: Piled embankment


Soil arching
Finite elements

1 Introduction In budget 2019, the Australian Government has committed to invest more than 75B AUD in public transport infrastructure projects over the next decade. Most of these projects needs to be constructed in the densely populated coastal regions. However, the presence of soft soil in the coastal regions poses a severe challenge to the stability of these infrastructure projects. Various ground improvement techniques have been used in the past to mitigate the issues associated with soft soil such as stage construction, deep foundation and preloading (Han and Gabr 2002; Bergado and Teerawattanasuk 2008). Columns (i.e., deep foundations) are considered a reliable technique for soft soil improvement as it reinforced to soft soil (Nunez et al. 2013; Yi et al. 2017). Recently, embankment with rigid columns, such as piles is used in practice due to its ability in reduced construction time and reduced displacement (Zhang et al. 2016). In the piled embankment, three components namely (i) embankment fill, (ii) load transfer platform, and (iii) rigid piles are responsible for surface and embankment load transfer to stiff stratum. However, some additional components such as geosynthetics and pile caps can also improve load transfer mechanism. Majority of imposed vertical © Springer Nature Switzerland AG 2020 P. Pinto et al. (Eds.): GeoMEast 2019, SUCI, pp. 40–50, 2020. https://doi.org/10.1007/978-3-030-34190-9_4

Finite Element Modeling of Soil Arching in Pile Supported Embankment

41

stress is transferred to the piles (due to their higher stiffness) by arching while a limited amount of stress is transferred to the subsoil (Terzaghi 1943). Several numerical as well as experimental studies have been carried out to investigate the arching mechanism in a piled embankment (Jenck et al. 2007; Potts and Zdravkovic 2008; Briançon and Simon 2012; Fagundes et al. 2015). However, these studies did not capture the arching mechanism under transport environment. In this study, the finite element method (FEM) is used to provide an insight into the behavior of the piled embankment with emphasis on the soil arching.

2 Numerical Modeling The FEM based software ABAQUS, version 6.13 is used to analyze a piled embankment in two-dimensional (2D) plane-strain condition. The 2D model is much easier and less time consuming without need of much advanced computing resources that would otherwise require for 3D modeling. Several idealization methods to convert 3D to 2D state are available (Bergado and Teerawattanasuk 2008; Huang and Han 2009). In this study, an Equivalent Area (EA) method is adopted for 3D idealization to 2D plane-strain. A unit cell is modeled to simulate the soil arching in a piled embankment. The modeled unit cell is assumed at the center of the piled embankment, passing through the half of the pile at both sides. Figure 1 shows a typical schematic diagram of modeled unit cell for a piled embankment with embankment height (he) 5.0 m and center to center pile spacing (s) 2.5 m. Both vertical sides of the unit cell represent lines passing through the centerto-center pile spacing (s) and overlying the embankment fill, including the gravel bed. The vertical boundaries are considered as roller supports restricting lateral movements, while the bottom boundary is fully fixed. The embankment top can move freely along vertical and lateral directions. The pile diameter (d) is assumed to be 1.0 m, whereas the center-to-center pile spacing (s) is considered to vary from 2.0 to 3.5 m. Different embankment heights (he) are considered. A 400 mm thick gravel bed is considered at the embankment base. The length of the pile and depth of the subsoil are considered the same, i.e. 8 m. A linearelastic, perfectly plastic model with Mohr-Coulomb failure criterion is considered for the embankment and subsoil material. The material parameters for piled embankment model are listed in Table 1.

3 Results and Discussion 3.1

Soil Arching

The variation in vertical stress (r) and settlement (d) above point A and B is referred to analyze soil arching (Fig. 1). The stresses, embankment heights, settlements are normalized for the sake of general applicability. Figure 2 shows the variation of the normalized vertical stress (rs/cs) against the normalized embankment height (he/ s) above point A for different embankment heights. Initially, the vertical stress increases

42

N. K. Meena et al.

Fig. 1. Typical schematic diagram of modeled unit cell Table 1. Material properties used for finite element simulation Material

Pile

Embankment fill Gravel bed Subsoil

20

Young’s modulus, E (MPa) 5000, 15,000, 30,000 5, 10, 30

21 18

65 40

Unit weight, c (kN/m3) 24

Poisson’s ratio, v

Cohesion c′(kPa)

0.15

–

0.25

0.1

0.25 0.3

0.1 10

Friction angle, /′ (degree) –

Dilation angle, w (degree) –

25, 30, 35, 40 35 15

20 20 0

Finite Element Modeling of Soil Arching in Pile Supported Embankment

43

with a gradient to geostatic stress from embankment top to the upper horizontal dotted line. The vertical stress show contrast trend, i.e. decrease in the zone bounded by the upper and lower horizontal dotted lines. The majority of the embankment load is transferred to the pile in this zone. Below the lower dotted line, vertical stress again shows an increase due to the soil self-weight. This variation in vertical stress along with embankment height is in good agreement with (Yu et al. 2016; Li et al. 2018). The soil arching is fully developed at 0.9he/s as the embankment height (he) exceeds 5.0 m. The lower boundary of arching is almost same for all embankment height, which implies its insensitivity to the embankment height while the outer boundary is primarily responsible for stress transfer to the pile head.

Fig. 2. Vertical stress variation in embankment fill over the subsoil (point A)

Figure 3 shows the variation of the normalized embankment settlement (de/s) with the normalized embankment height (he/s) above points A and B. The horizontal dotted line shows the plane of equal settlement. Above this plane, the embankment settlement does not change on points A and B, thus indicating the non-occurrence of the differential settlement. However, below this plane, the settlement over point A increases, while it decreases to almost zero over point B. It may thus be concluded that the soil

44

N. K. Meena et al.

Fig. 3. Settlement variation in embankment fill over the subsoil (point A) and pile head (point B)

arching is developed below the plane of equal settlement. The trend of Fig. 3 is in good agreement with other studies (Jenck et al. 2007; Li et al. 2018). Thus, the modeled unit cell can mimic the soil arching mechanism in the piled embankment. 3.2

Stress Concentration Ratio

The ratio between vertical stresses on the pile (rp) to the subsoil (rs) is termed as stress concentration ratio (SCR). It is expressed as (Han and Gabr 2002): SCR ¼

rp rs

ð1Þ

The SCR can evaluate the degree of stress concentration from soil to pile. SCR = 1 represents no soil arching (i.e., equal vertical stress on the pile head and subsoil) whereas a higher value of SCR implies to soil arching. Figure 4 shows the influence of embankment height (he) on the SCR. The SCR increases with increase in the embankment height until the maximum value is attained. For an embankment height (he) 2.5 m, maximum SCR is observed as 0.35he/s. It shows decrease until becoming

Finite Element Modeling of Soil Arching in Pile Supported Embankment

45

Fig. 4. Influence of embankment height (he) on stress concentration ratio (SCR)

unity if embankment height (he) continue to increase. For fully developed soil arching (i.e., he = 5.0 m), SCR becomes 0.20he/s. The embankment height associated with SCR = 1 implies to the upper boundary of the soil arching. It is observed that the upper boundary occurs at 0.9he/s. After this height, SCR remains constant, which implies that there is no stress redistribution. The soil arching boundaries are also confirmed in Fig. 2. Figures 5, 6 and 7 show the influence of piled embankment properties on stress concentration ratio. The influence of pile modulus (Ep) on the stress concentration ratio (SCR) is found to be negligible for all considered pile spacing (Fig. 5). The SCR was reported to increase with an increase of pile modulus (Han and Gabr 2002). However, in the current study, such a trend is not observed due to adoption of higher pile modulus. Figure 6 shows the effect of embankment modulus (Eem) on SCR. It is shown that SCR increases with an increase in Eem. However, the effect of embankment

46

N. K. Meena et al.

Fig. 5. Influence of pile modulus (Ep) on stress concentration ratio (SCR)

modulus is found to be negligible after a specific embankment modulus (i.e., Eem = 30 MPa). The effect of friction angle (/′) on the stress concentration ratio (SCR) is shown in Fig. 7. It shows that SCR increases by 40% as the friction angle is varied from 25° to 40°. Figures 5, 6 and 7 also show that the SCR decreases with an increase of pile spacing (s). The arching zone shows more pronounced mobilization even with a smaller reduction in s. Thus, the pile spacing (s) should be kept as minimum as possible to derive a higher degree of soil arching. 3.3

Shape of Soil Arching

In the past, the different soil-arch shapes are considered to develop the analytical models for piled embankment. Some of them are adopted in different design guidelines such as Dutch guideline (Van Eekelen et al. 2010), German guideline (EBGEO 2011) and French guideline (ASIRI 2012). However, a wide variation in results occurs due to

Finite Element Modeling of Soil Arching in Pile Supported Embankment

Fig. 6. Influence of embankment modulus (Eem) on stress concentration ratio (SCR)

Fig. 7. Influence of friction angle (/′) on stress concentration ratio (SCR)

47

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N. K. Meena et al.

different recommendations by each guideline even for the same condition. A more consistent approach is therefore needed. In view of this, the shape of soil arching is also investigated in this study.

Fig. 8. Shape of the soil arching

Figure 8 shows the shape of soil arching. It is evident that the thickness of soil arching is not uniform for all embankment height. The soil arch is thicker over point A (i.e., subsoil) compared to point B (i.e., pile head). Thus, one may conclude that the soil arch is not semicircular with uniform thickness and does not follow Hewlett and Randolph (1988). It is found to be identical to the multi-shell theory proposed earlier by Kempfert et al. (2004).

4 Conclusion A numerical investigation of soil arching in piled embankment under the plane strain condition has been reported in this paper. The effect of piled embankment parameters is highlighted. Based on the results, the following conclusions are made: • The arching mechanism is affected by piled-embankment properties. Pile and embankment modulli and friction angle significantly affect the soil arching.

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49

• The arching zone becomes more pronounced even with a smaller reduction in the pile spacing. • The arching zone is found to the maximum at the mid of subsoil whereas minimum on the pile head, which is identical to the multi-shell theory. Acknowledgments. The authors gratefully acknowledge the financial support provided by the Government of India under the National Overseas Scholarship, No. 11016/16/2016 Education.

References ASIRI P.: Recommendations for the design, design, execution and control of the improvement of foundation soils by rigid inclusions. Presses des Ponts (2012). (in French) Bergado, D.T., Teerawattanasuk, C.: 2D and 3D numerical simulations of reinforced embankments on soft ground. Geotext. Geomembr. 26(1), 39–55 (2008) Bhasi, A., Rajagopal, K.: Numerical investigation of time dependant behaviour of geosynthetic reinforcement piled embankment. Int. J. Geotech. Eng. 7(3), 232–240 (2013) Briançon, L., Simon, B.: Performance of pile-supported embankment over soft soil: full-scale experiment. J. Geotech. Geoenviron. Eng 138(4), 551–561 (2012) EBGEO: Recommendations for design and analysis of earth structures using geosynthetic reinforcements–EBGEO. Digital in English (2011) Fagundes, D.D.F., Almeida, M.D.S.D., Girout, R., Blanc, M., Thorel, L.: Behaviour of piled embankment without reinforcement. Proc. Inst. Civ. Eng. Geotech. Eng. 168(GE6), 514–525 (2015) Han, J., Gabr, M.A.: Numerical analysis of geosynthetic-reinforced and pile-supported earth platforms over soft soil. J. Geotech. Geoenviron. Eng. 128(1), 44–53 (2002) Hewlett, W.J., Randolph, M.F.: Analysis of piled embankments. Ground Eng. 21(3), 12–18 (1988) Huang, J., Han, J.: 3D coupled mechanical and hydraulic modeling of a geosynthetic-reinforced deep mixed column-supported embankment. Geotext. Geomembr. 27(4), 272–280 (2009) Jenck, O., Dias, D., Kastner, R.: Two-dimensional physical and numerical modeling of a pilesupported earth platform over soft soil. J. Geotech. Geoenviron. Eng. 133(3), 295–305 (2007) Kempfert, H.-G., Göbel, C., Alexiew, D., Heitz, C.: German recommendations for reinforced embankments on pile-similar elements. In: Proceedings of EuroGeo3, Munich, pp. 279–284 (2004) Li, X., Miao, Y., Cheng, K.: Soil arching effect analysis via a modified finite element model based on a field test. J. Test. Eval. 46(5), 2218–2226 (2018) Nunez, M.A., Briançon, L., Dias, D.: Analyses of a pile-supported embankment over soft clay: full-scale experiment, analytical and numerical approaches. Eng. Geol. 153, 53–67 (2013) Potts, V.J., Zdravkovic, L.: Finite element analysis of arching behaviour in soils. In: 12th International Conference on Computer Methods and Advances in Geomechanics, pp. 3642– 3649 (2008) Terzaghi, K.: Theoretical Soil Mechanics. Wiley, New York (1943) Van Eekelen, S.J.M., Jansen, H.L., Van Duijnen, P.G., De Kant, M., Van Dalen, J.H., Brugman, M.H.A., Van, D.S., Peters, M.G.J.M.: The Dutch design guideline for piled embankments. In: 9th International Conference on Geosynthetics, Geosynthetics for a Challenging World, Guaruja, Brasil, pp. 1911–1916. IGS (2010)

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Yi, Y., Liu, S., Puppala, A.J., Xi, P.: Vertical bearing capacity behaviour of single T-shaped soil– cement column in soft ground: laboratory modelling, field test, and calculation. Acta Geotech. 12(5), 1077–1088 (2017) Yu, Y., Bathurst, R.J., Damians, I.P.: Modified unit cell approach for modelling geosyntheticreinforced column-supported embankments. Geotext. Geomembr. 44(3), 332–343 (2016) Zhang, C., Jiang, G., Liu, X., Buzzi, O.: Arching in geogrid-reinforced pile supported embankments over silty clay of medium compressibility: field data and analytical solution. Comput. Geotech. 77, 11–25 (2016)

Study of Slope Stability Using Piles, Pathological Case in Mostaganem, Algeria A. Belghit, S. M. A. Bourdim(&), A. Benanane, and N. Bouhamou LMPC Laboratory, Department of Civil Engineering and Architecture, University Abdelhamid Ibn Badis of Mostaganem, Mostaganem, Algeria [email protected]

Abstract. The construction on a site characterized by a slope is possible, but exposed to several geotechnical phenomena (particularly the landslides). The latter can cause real problems with the stability of structures if the necessary precautions will not be taken in consideration by the designer. The objective of our work is the study of the stability of the slopes by the piles where we will present the risks of instability, the methods of calculation of the safety coefficient and determine the effective position of the piles that allows to improve the security of unstable slopes using a numerical calculation by the code Plaxis 2D. This study will be followed by an application of a pathological case in the field of Public Works. Keywords: Landslide


Stabilization
Anti-slidings piles
Safety coefficient

1 Introduction The study of sloped terrains has always been the center of interest of many specialists and experts in civil engineering due to their constant exposition to natural, geological and hydraulic phenomena (earthquakes, floods, erosions, landslides…etc.). The main issue that caused the instability of sloping sites are the landslides that damaged villages, roads, dams and natural slopes. The origins of the loss of stability, both natural and artificial, are very diverse. The setting in motion of natural slopes (slow or sudden) may cause significant damage to structures, with a significant economic impact, and sometimes injury or death to humans. The study of a slope comprises, besides the recognition of the site and the selection of the mechanical characteristics of the soil, a stability calculation to determine firstly the failure curve along which the risk of slip is highest, the surface along which the factor of safety is the lowest. Landslides are diverse, both in their nature and size. The landslide passes through several chronological stages of activity. There are main factors that control the type and rate of mass movements that might occur [5]. Several researches have been developed to increase the slope stability by different methods as the earthworks protections, inclusions…etc.

© Springer Nature Switzerland AG 2020 P. Pinto et al. (Eds.): GeoMEast 2019, SUCI, pp. 51–60, 2020. https://doi.org/10.1007/978-3-030-34190-9_5

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In this research, we will analyze a pathological case occurred at Mostaganem in Algeria. And we will try to apply our solution to solve this type of problems.

2 Presentation of the Pathological Case In this article we discuss the landslide which occurred on December 2016, located in the national road RN90 just at PK58+200 next to the village of TEZGAIT in the limit between Mostaganem city and Relizane city (the coordinates of the site location N36 09478 E 0 552970) as it is shown in the next satellite photo. We visited the landslide site in 20/03/2017 to provide a decision on the damage caused by the land movement in January 2017 (Fig. 1).

Fig. 1.

Satellite view of the landslide site

The probable cause of this situation is water and exactly the rainwater of December 2016. The slip plan was formed for several years due to many cracks that have been noticed on site in 2015. The maintenance efforts could not stop the movement process of this area until the break in January 2017 (Fig. 2). The probability for a landslide to happen again in this site is not expected because of the absence of other cracks on the paving according to our visual observation. The following photos was taken on site:

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Photo 1.

53

Photo taken at the unstable site

3 Geotechnical Parameters These information were taken from a geotechnical document at the administration. 0–0.35 m: topsoil, 0.35–1 m: dark brownish clay and 1 m–5 m: greyish marne concreted. No trace of water appeared. Table 1 show the geotechnical parameters of the site.

Table 1. Geotechnical parameters of soil layers Friction angle / (°) 30

Cohesion C (kPa) 5

E (kPa) 2000

#

17

csat (kN/m3) 18,33

19

19,63

15

20

2000

0,35

Soil

c (kN/m3)

Dark brownish clay Greyish marne concreted

0,35

And to modeling the concrete material’s for piles we use the elastic model provided by the PLAXIS software.

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Fig. 2. Topographical survey of the landslide site.

4 Topography of the Site With the help of the research in the archives of the public works department we were able to have the topographical survey of the site that experienced the landslide as shown in this figure.

5 Finite Elements Model of the Landslide To determinate the sensitive places that risk destabilization or breakage, in the first place we will proceed to the modeling of the cross section of our field to determine the safety factor of our case and to determine the sensitive points. The following figure shows the modeling of our soil in PLAXIS 8.2 software over a model width of 70 m and a height of 70 m with the real dimensions of the profile across the slope of our case (Figs. 3 and 4).

Fig. 3.

Cross section of ground at sliding level.

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Fig. 4. Finite elements model of the site

After we looked over the results obtained during the modeling of our site using Plaxis 2D software, we noticed that the embankment located on the upper part of the road is unstable which applies an additional constraints on the lower part of the road and that will put the roadway in permanent danger which explains the movement of our land (Fig. 5).

Fig. 5.

Critical slip surface

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The safety factor equals to 0.895 (less than 1): this value explains that the ground is in a state of rupture.

6 Treatment of the Pathology According to the results obtained in the previous part, from where a distance of 3/4 L represents the best stability of the slope, the pile was modeled at this position (Figs. 6 and 7).

Fig. 6. Mesh of deformation due to self weight of the slope. Safety factor = 0.89.

Fig. 7. Stabilization of the site by two piles.

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This safety factor value (1.46) shows a slope in a state of stability but remains a value below 1.50 (safety factor value in safety state) but it can be improved, for example to improve this value it suffices to minimize the motor loads, in particular the reduction of the weight of soil at the top of the embankment by the execution of disbursements (Figs. 8 and 9).

Fig. 8. Deformation mesh after introduction of the piles.

Fig. 9. Failure mechanism after introduction of the piles. Safety factor = 1.46

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In order to ensure that the safety factor value will be improved by lightening the top of the embankment, we will proceed to the modeling of our case with the disbursement in the upper embankment bench implanting a pile in 3/4 L of the part of the slope remaining to see the improvement of the security of our case (Figs. 10 and 11).

Fig. 10.

Fig. 11.

Slope modeling with piles and platform.

Mesh deformation of slop with piles and platform.

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7 Discussion The results obtained in this study that the positioning of a pile wire at ¾ L (L: transverse width of an embankment) to stabilize a slope that risks breaking gives it a good safety and can be generalized for most cases of slope stability (Fig. 12).

Fig. 12.

Failure mechanism of the slop with piles and platform. Safety factor = 1.74

8 Conclusion After developing this study (stability of slopes by piles) we can say that the stabilization of slopes by piles is a method that has a very good reliability for the range of slopes of medium and large dimensions. So for cases that cannot be mastered by conventional and simple maintenance methods (earthworks, drains and purges…etc.) the second range is stabilization by rigid inclusions and especially the piles is preferable. In addition, the realization of piles by threshing or drilling (on-site concreting) for local companies are now becoming very practical and simple in terms of financial equipment or duration of implementation. At least one pile wire is required for an embankment that presents a risk of rupture or in a state of motion or as a prevention for cases where the factor of safety is less than 1.3. These piles should be implanted near the ¼ L top of slope (L the transverse slope width). The numerical calculation of the factor of safety must be carried out for the two cases of the same slope to be stabilized without and with piles in order to see the improvement after inclusion and to ensure a coefficient of safety higher than 1.5. At the end, we can say that this work has allowed us to chart a new path in this type of study and hoping that it will be a starting point in scientific research for our working group and for future generations.

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References 1. Durville, J., Seve, G.: Stabilité des pentes- Glissements en terrain meuble. Tech. l’ingénieur (1996) 2. Bousquet, P.: Pieux et palplanches, Tech. l’lngénieur, pp. 1–32 3. Ouzaid, I.: THEME : Stabilisation des Pentes par Inclusions Rigides (2014) 4. Chabil, H., Baheddi, M., Hamoud, F., Karech, T., Abbeche, K.: Stabilisation des talus renforces par pieux 5. Bourdim, S.M.E.-A., El-Hakim Chekroun, L., Benanane, A., Bourdim, A.: Treatment of a Landslide by Using Piles System, Case Study of the East-West Highway of Algeria (2018) 6. M. E. M. K. ép. Sekkel: Cours De Mecanique Des Sols 1 7. Fırat, S.: Stability analysis of pile-slope system. Sci. Res. Essay 4, 842–852 (2009) 8. Liu, Z.: Evaluation et amélioration des modèles numériques pour l’analyse de la stabilité des pentes Soutenue (2013) 9. Khemissa, M.: Méthodes d’analyse de la stabilité et techniques de stabilisation des pentes. Géologie Géotechnique l’Ingénieur JNGG (2006) 10. Lefriki souad: Effet De La Variation De La Cohetion Sur Le Comportement Des Pentes (2015) 11. Fontaine, M., Genivar, W.S.P.: Stabilisation de pentes 12. Munawir, A., Dewi, S.M., Zaika, Y., Soehardjono, A.: Safety factor of continuous footing on slope modeling with composite bamboo pile reinforcement. Electron. J. Geotech. Eng. 18, 2177–2186 (2013)

Effect of Fines and Matric Suction on the Collapsibility of Sandy Soils Mohamed A. Alassal(&), Asmaa M. Hassan(&), and Hussein H. Elmamlouk(&) Cairo University, Giza, Egypt [email protected], [email protected], [email protected]

Abstract. New geotechnical challenges are encountered due to construction of new urban communities in desert, among which is dealing with collapsible soil formations. Collapsible soils are metastable material, traditionally defined as an unsaturated soil that experiences a radical rearrangement of particles and significant reduction of volume upon wetting. In this study, an experimental program is conducted to investigate the influence of various parameters on the collapse of reconstituted sandy soils. The magnitude of collapse of ten sandy soils containing different types and percentages of fines is determined using the single oedometer Test. All these soils are prepared at 35% relative density. Then, the effect of related parameters including silt content (10%–50%), type of fines (silt/clay), initial water content (5%–15%), and wetting pressure (0– 200 kPa) have been studied. Furthermore, the initial matric suction is determined using the ASTM filter paper method in order to study its effect on the soil collapsibility. Finally, Scanning Electron Microscope (SEM) is used to visually examine the effect of fines on the collapse susceptibility of sandy soils.

1 Introduction Geotechnical engineers encounter immense challenges while dealing with collapsible soils as urban communities are expanding towards desert. Collapsible soils cover approximately 10% of the world land (Jefferson et al. 2008) and are reported to exist in several areas of Egypt. However, most studies focus on new urban developments such as 6th of October, Sheikh Zayed, 10th of Ramadan, Nasr City, New Maady, West Cairo, New Amerya, and Sadat City (Mossaad et al. 2008). Collapsible soil is defined as an unsaturated soil which sustains high values of applied pressures at their natural water content. Upon wetting, this type of soil experiences a significant and sudden reduction in volume. The lack of understanding of the behavior of collapsible soils has caused a lot of damages to infrastructures, buildings and other structures (Mossaad et al. 2008). According to previous research, the magnitude of collapse depends on various factors among which; initial water content, initial dry unit weight, soil gradation, and wetting pressure (Basma and Tuncer 1992; Ayadat and Hanna 2007; Phanikumar et al. 2015). Attempts have been made to use constitutive relations for unsaturated soils to predict the volume change of a collapsible soil during inundation. It is found that the © Springer Nature Switzerland AG 2020 P. Pinto et al. (Eds.): GeoMEast 2019, SUCI, pp. 61–72, 2020. https://doi.org/10.1007/978-3-030-34190-9_6

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collapse phenomenon is primarily related to the reduction of the matric suction during inundation (Ramos and Valencia 2013). Matric suction is one of the two stress-state variables that control the behavior of an unsaturated soil. Matric suction is attributed to soil texture, capillary and adsorptive forces acting between liquid, air, and solids. Matric Suction could be measured by direct (e.g. tensiometer) and indirect (e.g. filter paper) methods (Fredlund and Rahardjo 1993). This paper presents the results of an experimental program conducted on ten sandy soils using the single oedometer test and filter paper method. In addition, index property tests are done including sieve analysis, hydrometer, Atterberg limits, and maximum and minimum void ratio. The current study examines the effect of related parameters including silt content, type of fines, initial water content, wetting pressure on the soil collapsibility. Moreover, the impact of variations in matric suction on the magnitude of collapse has been investigated. In addition, visual inspection of the micro-structure of selected soil samples is performed using scanning electronic microscope (SEM) images.

2 Materials and Tests Ten collapsible sandy soils are prepared in the laboratory from a mixture of sand, nonplastic silt, and clay. Fine to medium siliceous sand from New Giza, Cairo is used in this study, whilst, the non-plastic silt is obtained by grinding this sand. The clay is collected from a quarry located in Fostat city, Cairo. A series of laboratory tests is conducted on the raw materials and the prepared collapsible soils including sieve analysis and hydrometer test, Atterberg limits, specific gravity, single oedometer collapse test, and filter paper method. According to USCS, the sand is classified as SP (poorly graded sand). The coefficient of curvature (cc) is 0.84 and the coefficient of uniformity (cu) is 3.67. For nonplastic silt, the coefficient of curvature (cc) is 2.84 and the coefficient of uniformity (cu) is 18.50. The grain size distribution curves are shown in Fig. 1a. The Atterberg limits tests are also determined. For clay, the liquid limit (wL), plastic limit (wP), and plasticity index (Ip) are 88%, 29% and 59%, respectively. Hence, clay is classified as CH (high plastic clay). For silt, the liquid limit (wL) is 28%, however, plastic limit could not be measured. Hence, silt is classified as ML (non-plastic silt). The specific gravity of the three soils equals 2.64. The collapsible sandy soils are divided into two groups. Group (A) is designated to study the effect of silt content. In this group, the soil mixture is composed of 10% clay and varying percentages of sand and silt (Fig. 1b). On the other hand, Group (B) is designated to mainly study the effect of fines type. In this group, each soil mixture is composed of 60% sand and 40% fines. Several combinations of silt and clay are used to prepare different types of fines (Fig. 1c). For both groups, all the soils are prepared at a relative density of 35%. Hence, for each soil type, maximum and minimum void ratios (emax, emin) are determined. The physical properties of these soils are provided in Table 1. The procedure adopted to prepare the specimens is as follows: First, the required percentages of sand, silt, and clay are mixed. Second, the dry weight required to

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Fig. 1. Grain size distribution curves (a) raw materials (b) Group (A) soils and (c) Group (B) soils

Group Raw material Soil mixture GSDC code Sand Silt content Clay D60 (mm) content (%) (%) content (%) A 50 40 10 S50-M400.110 C10 70 20 10 S70-M200.230 C10 80 10 10 S80-M100.270 C10 90 0 10 S90-M000.300 C10 A,B 60 30 10 S60-M300.009 C10 B 60 40 0 S60-M400.180 C00 60 20 20 S60-M200.034 C20 60 10 30 S60-M100.180 C30 60 2 38 S60-M020.180 C38 0.180 60 0 40 S60-M00C40 *Np is not applicable and could not be estimated for this soil type 0.180 0.011 -

0.045 0.053 0.160 0.018 0.018 -

-

-

-

16.4

20.0

4.0

-

-

-

-

0.075

0.115

-

0.006

0.095

88

86

79

54

1.42 28

1.25 42

0.60 88

5.60 54

115.0 12.20 48

0.002

0.075 45.0

-

-

Fines LL (%) 35

Cc

Cu

properties D30 D10 (mm) (mm) 0.030 -

Table 1 Physical Properties of collapsible soils (Dr = 35%)

24

59

35

30

29

29

27

19

59

57

52

35

NP* -

18

29

19

18

properties PL PI (%) (%) 15 20

SC

SC

SC

SC

SM

SPSC SC

SC

SC

SC

1.21 0.75 1.05

1.20 0.60 0.99

1.09 0.49 0.88

1.05 0.45 0.84

1.02 0.42 0.81

1.00 0.43 0.8

0.90 0.47 0.75

0.81 0.38 0.66

0.99 0.39 0.78

1.10 0.47 0.88

USCS Void ratio emax emin eused

64 M. A. Alassal et al.

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achieve a specified void ratio value corresponding to a relative density of 35% is calculated. Third, a calculated amount of distilled water is added to soil in order to reach the required initial water content. The moist soil is thoroughly mixed and kneaded by hand to form a homogenous mixture. Finally, specimens are prepared by placing the moist soil in layers into a mould (53 mm in diameter and 52 mm in height). Each layer is tamped using a designed hammer to the required height corresponding to the specified void ratio. Single oedometer collapse test (ASTM D5333, 03) and filter paper method (ASTM D5298, 03) are performed to evaluate the soil collapsibility and matric suction. Single oedometer test is used to determine the collapse potential. After preparing specimens, as previously explained, the cutting edge of the oedometer ring is centered at the top of specimen and is slowly pushed down while trimming the soil around the ring. After fully penetrating the sample, soil above and below ring is trimmed. Then, the load is applied gradually up to the designated wetting pressure, at which the specimen is flooded with distilled water and left for 24 h. Dial gauges are mounted on the top of the specimen for measuring the vertical displacement. The collapse potential is calculated using the following equation: CP =

De 1 + eo

Where De is the change in void ratio due to inundation and eo is the initial void ratio. The filter paper technique is adopted to measure the soil matric suction using filter paper disks (Whatman No. 42, ashless). This method is adopted as a simple and feasible alternative which can reliably measure a wide range of suction values (Fredlund and Rahardjo 1993). This test involves preparing a pair of nominally identical specimens. Next, a stack of three filter papers (composed of a filter paper disk placed between two protective filter papers) is inserted between the two identical specimens. To prevent soil contamination, the filter paper disk is sandwiched between two protective filter papers that have a slightly larger diameter (5 mm larger) than the central disk. The specimens are immediately sealed with a plastic sealing tape and wrapped with foil paper. To enhance the contact between the filter paper disk and the soil, a vertical stress of 5 kPa is applied. The direct contact between the filter paper and the soil allows water in the liquid phase and solutes to exchange freely (Fredlund and Rahardjo 1993). The specimens are kept in a desiccator for 7 days. During this period, water content of the initially dry filter paper increases until equilibrium is established between filter paper and soil. Then, the water content of the filter paper is measured. Finally, using the appropriate filter paper calibration curve, soil matric suction is estimated (Kim et al. 2016).

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3 Results and Discussions Figure 2 shows the typical relationship obtained from a single oedometer collapse test. It presents the change in void ratio (e) as the applied pressure (P) increases. This e-P relationship is divided into three segments; before wetting, at wetting, and after wetting which presents the typical behavior of a collapsible soil formation. In the presence of fines acting as a binding agent and large matric suction values, soil exerts small values of compression despite of its low initial relative density (Dr = 35%). Upon wetting, as the water content increases, the binding is destroyed and the matric suction is reduced, consequently, the soil collapses (Mitchell 1976). The third segment illustrates the behavior of collapsible soil after inundation. Figure 3a investigates the effect of silt content on the soil collapsibility using Group (A) soils while maintaining a clay content of 10% and an initial water content of 5%. Despite the fact that all specimens are prepared at the same relative density, it is noted that as silt content increases, collapse potential increases. If no silt is added to the prepared soil mixture, collapse potential is 3.3%. If a silt content of 40% is used, collapse potential increases to 11.7% (255% increase) which agrees with the trend observed by Steadman (1987). This trend is comparable to the output deduced from Fig. 3b. Soil with a higher percentage of silt content has a higher initial matric suction. Moreover, as silt content increases, a reduction in the rate of increase of collapse is observed. Increasing silt content from 0% to 10% leads to an increase in collapse potential by 84%. However, increasing silt content from 30% to 40% leads to an increase in collapse potential by 6% only (Fig. 3a). The same trend is observed in Fig. 3b, the rate of change in the measured matric suctions decreases as silt content increases. This phenomenon can be explained based on the work of Ni et al. (2004). The non-plastic silt particles may not act as a binding agent, therefore, the space they occupy may be considered as voids. Hence, the skeleton void ratio (es) is proposed to be used instead of the void ratio (e). es is defined as the void ratio of soil if fines are considered as voids rather than solid particles. As shown in Fig. 4, as the silt content increases, the skeleton void ratio (es) increases, consequently, collapse potential (CP) increases. The effect of fines type is studied using Group (B) soils. Six mixtures are tested with different types of fines. These mixtures have different percentages of silt and clay, however, the total percentage of fines is maintained at 40%. In order to establish a relationship between fines type and collapse potential (CP), fines type is expressed by its plasticity index (PI). The plasticity index (PI) increases upon using higher clay contents. For clay content = zero and 40%, PI = 0 and 59%, respectively. Soils containing higher clay contents (higher PI values) experience higher magnitudes of collapse (Fig. 5a). The effect of fines type is also reflected on the measured values of initial matric suction. The initial matric suction is 135 kPa, if no clay is added (PI = 0), however, it reaches a value of 36,000 kPa if clay content of 40% is used (PI = 59%) as observed in Fig. 5b. Clayey fines between sand grains occurs in the form of several shapes (Alwail 1990) among which is “clay balls”. SEM micrograph in Fig. 6 shows clay balls in two different soils. Most of these balls disintegrate upon inundation, consequently, reduction in volume takes place. It is noticed that for clay content of

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30%, bigger clay balls are formed compared to soil with clay content of 10%. The development of bigger clay balls is associated with the formation of larger void spaces, consequently, higher values of collapse upon inundation. The initial water content (wi) has a significant effect on the measured collapse potential and initial matric suction (Fig. 7). For soil mixture S60-M30-C10 at wi = 2.5%, CP = 21%. If wi increases to 15%, CP decreases to a value of 1.5% (93% decrease) as observed in Fig. 7a. As wi increases, lower values of initial matric suction are induced and weaker bonds provided by fines are formed, thus, the soil microstructure becomes unstable (Fig. 7b). Consequently, higher values of compression are depicted under applied pressure even at the initial water content (before inundation). This conclusion is supported by Basma and Tuncer (1992) and Phanikumar et al. (2015).

Fig. 2. Single Oedometer collapse test results: void ratio versus load

Figure 8 shows that collapse potential (CP) increases with increasing wetting pressure (Pw). However, the increasing rate of CP declines as Pw increases until Pw = 100 kPa. Beyond this value, CP remains almost constant or slightly decreases. The results agrees with the work of Rabbi et al. (2014). This can be attributed to two reasons. First, there is a maximum degree of densification which can be triggered by the collapse mechanism and is achieved at a certain pressure level. Any further increase in pressure beyond this level will, subsequently, cause little or no change in collapse potential. Second, applying high levels of pressure leads to the partial breakage of the binding agents before inundation.

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Fig. 3. Effect of silt content on (a) Collapse potential and (b) soil matric suction

Fig. 4 Skeleton void ratio (es) vs collapse potential (CP) (Dr = 35% - wi = 5%)

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Fig. 5 Effect of fines type on (a) collapse potential and (b) soil matric suction

(a)

(b)

4 3 2 2 1 1 3 4

Fig. 6 SEM showing micro-structure of collapsible soil specimens at dry condition [1] sand particle, [2] void, [3] clay balls, [4] silt for (a) S60-M30-C10 and (b) S60-M10-C30

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Fig. 7 Effect of initial water content on (a) collapse potential (CP) and (b)

Fig. 8 Effect of wetting pressure (Pw) on Collapse potential (CP)

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4 Conclusions This paper presents the results of an experimental work conducted on ten reconstituted sansy soils. The effect of silt content, fines type, initial water content, and wetting pressure on the collapse potential is investigated. It is found that the non-plastic silt particles act as voids rather than binding agents. Therefore, increasing non-plastic silt content (while sustaining the same clay content) leads to increase in the total volume of voids (voids + space occupied by silt), accordingly, collapse potential increases. The study also reveals that collapse potential increases with increasing clay content up to 40%. In this case, the dominant mechanism responsible for collapse (upon inundation) is the disintegration of clay balls. Using a higher value of initial water content weakens the soil micro-structure which leads to a larger magnitude of collapse. For the tested soils, collapse potential increases with increasing wetting pressure up to a certain value of Pw, after which collapse potential remains the same or decreases with further increase in wetting pressure. In addition, the filter paper method is adopted as a feasible technique to measure the soil matric suction. The variation in the measured suction values are found to be directly influenced by the change in the soil related properties which affect the magnitude of collapse. Therefore, the soil matric suction can be used as a key parameter to predict the soil collapsibility.

References Alwail, T.A.: Mechanism and effect of fines on the collapse of compacted sandy soils. Doctoral dissertation, Washington State University, Washington (1990) American Society for Testing: Standard test method for measurement of soil potential (suction) using filter paper. 2003 Annual Book of ASTM Standards, vol. 04.08, D 5298-03. ASTM (2003) American Society for Testing: Standard test methods for measurement of collapse potentials of soils. 2000 Annual Book of ASTM Standards, vol. 04.08, pp. 343–345, D5333–03, ASTM, Philadelphia (2003) Ayadat, T., Hanna, A.: Prediction of collapse behaviour in soil. Revue Européenne de Génie Civil 11(5), 603–619 (2007) Basma, A.A., Tuncer, E.R.. Evaluation and control of collapsible soils. J. Geotech. Eng. Div. 115 (9), 1252–1267, 118, 1491–1504 (1992) Fredlund, D.G., Rahardgo, H.: Soil Mechanics for Unsaturated Soils. Wiley, New York (1993) Jefferson, I., et al.: The treatment of collapsible loess soils using cemented materials. GeoCongress 2008, 662–669 (2008) Kim, H., et al.: Calibration of Whatman grade 42 filter paper for soil suction measurement. Candian J. Soil Sci. 97(2), 93–98 (2016) Mitchell, J.K.: Fabric, structure, and property relationships. In: Fundamental of Soil Behavior, pp. 222–252. Wiley, New York (1976) Mossaad, E.M., et al.: A study on collapsing soils in Egypt. Internal Research (2008) Report, Academy for Scientific Research & Technology (ASRT), Egypt Ni, Q., et al.: Contribution of fines to the compressive strength of mixed soils. Geotechnique 54 (9), 561–569 (2004)

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Phanikumar, B.R., et al.: Collapse behaviour of lateritic soil. Int. J. Geotech. Eng. 11(2), 119–124 (2015) Rabbi, A.T., Cameron, D.A., Rahman M.M.: Role of matric suction on wetting induced collapse settlement of silty sand. In: Khalili, N., Russell, A., Khoshghalb, A. (eds.) Unsaturated Soils, Research & Applications, pp. 129–135 (2014) Ramos J., Valencia, Y.: Evaluation of soil matric suction, microstructure and its influence on collapsible behavior. In: Advances in Unsaturated Soils. Taylor & Francis Group, London (2013) Steadman, L.: Collapse settlement in compacted soils of variable fines content. Doctoral dissertation, Washington State University, Washington (1987)

Assessment of Earth Retaining Wall Sustainability: Value Functions and Stakeholder Weighting Sensitivity I. P. Damians1,2(&), R. J. Bathurst3, A. Lloret1, A. Josa1, and D. El Mansouri2 1

School of Civil Engineering, Universitat Politècnica de Catalunya BarcelonaTech, Barcelona, Spain [email protected] 2 VSL International Ltd., VSL Technical Center, Barcelona, Spain 3 Royal Military College of Canada, Kingston, ON, Canada

Abstract. Earth retaining walls are common geotechnical structures with a wide range of solutions available to perform the same function. More and more, geotechnical engineers are asked to find the best solution among several options in different civil engineering applications based on environmental impact, cost and societal/functional issues. Evaluation of these three pillars during the selection process of a structure (such as an earth retaining wall) is called a sustainability assessment. This paper describes a sustainability assessment methodology and gives examples to select the best sustainable option from candidate conventional gravity and cantilever wall types, and steel and polymeric soil reinforced mechanically stabilized earth (MSE) walls of 5 m height. Analyses were carried out using the MIVES methodology which is based on value theory and multi-attribute assumptions. The paper identifies how indicator issues are scored, weighted and aggregated to generate final numerical scores that allow solution options to be ranked. The final scores include an adjustment based on stakeholder preferences for the relative importance of the three sustainability pillars (environmental, economic (cost) and societal/functional). The analysis results show that MSE wall solutions are most often the best option in each category compared to conventional gravity and cantilever wall solutions and thus most often the final choice when scores from each pillar were aggregated to a final score. The paper also includes a sensitivity analysis of the choice of value functions and stakeholder weighting preferences on the final ranking scores used to select the best sustainable solution. The analyses also show that the choice of value function and stakeholder preferences can lead to a conventional structure being the best option.

1 Introduction Sustainability and sustainable development relate to the capacity to carry out an activity (such as manufacturing or constructing a product) with minimal economic, societal and/or environmental impact (WCED 1987; Josa and Alavedra 2006). In civil engineering works, several construction solutions can satisfy the same functional © Springer Nature Switzerland AG 2020 P. Pinto et al. (Eds.): GeoMEast 2019, SUCI, pp. 73–88, 2020. https://doi.org/10.1007/978-3-030-34190-9_7

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requirements. Examples are pavements (asphalt, concrete…), bridges (simple beam, arch, cable-stayed, suspension…), foundations (shallow, deep…), etc. However, solutions will vary with respect to sustainability (i.e., different costs of materials, construction and maintenance, different environmental impact and societal effects). The adoption of sustainability criteria during the design of new projects is becoming more common in many countries, and these criteria are being used to influence decision making (Aguado et al. 2012). In the past, the final selected solution was typically based on a compromise between minimum costs and maximum functionality. Today, an appropriate sustainability analysis approach is often recommended so that environmental and social impacts are part of the solution decision process. From a total sustainability point of view there are three pillars or requirements: environmental, economic and societal that can be assessed from “cradle to grave” (though in some situations “cradle to gate” or “cradle to operational” is more realistic). A proper balance between these criteria is required for sustainable development or solution selection (Josa and Alavedra 2006; Josa et al. 2008). Earth retaining walls are very common in civil engineering works with wellestablished design procedures for typical structures. Nevertheless, there are a wide range of solutions within any wall classification and between different wall solution classifications (e.g., conventional gravity, cantilever or mechanically stabilized earth (MSE) walls). A discussion of the advantages and limitations of different types of earth retaining wall solutions can be found in guidance documents and in papers such as Jones (2002) and Damians et al. (2017, 2018). Sustainability is a young discipline, and a consensus on a formal set of rules is not available, particularly in civil engineering. Different models continue to be developed or refined. This paper uses the case of earth retaining wall structures to make choices between wall options based on a sustainability perspective (Damians et al. 2018) including sensitivity analyses on value function types and stakeholder preferences. The paper identifies the information that is required to exercise rational decision making for these types of structures within a sustainability assessment framework.

2 Methodology 2.1

Sustainability Model: MIVES Tool

There is no unique way to quantify sustainability. Therefore, a comparison methodology is required with application-specific sustainability models. As a starting point, it is necessary to identify the same functional unit (FU). This unit defines what precisely is being studied and provides a reference for the inputs and outputs that enable the alternatives to be compared and analyzed. A sustainability model assessment based on value-theory and multi-attribute assumptions called MIVES (Value Integrated Model for Sustainable Evaluations; Josa et al. 2008) was adopted in this investigation to evaluate the sustainability of each case study. Each sustainability requirement (i.e., environmental, economic, societal/functional) can be defined by one or more criteria and a set of quantifiable indicators. The set of requirements, criteria and indicators define a decision-making tree for multi-criteria analysis. In other models, additional

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criteria levels are possible, but in this investigation a simple approach was adopted. The indicators are defined by value-functions which can have different forms (i.e., linear, concave, convex, S-shape; see Alarcon et al. 2011). The value-functions allow the transformation from indicator-units (e.g., physical units) to common dimensionless value units. Next, the related criteria and indicators are defined globally by weighting and aggregation procedures. The final result is an overall single index value (final score) for each proposed alternative. This process is a powerful tool when an objective decision is needed (i.e., as opposed to a purely subjective decision), and allows the weaknesses and strengths of each proposal to be assessed. Strategies to analyze the three pillars of sustainable development are briefly explained below. The environmental pillar can be defined using specific indicators applicable to each study case. A powerful but sometimes difficult approach is to develop a life cycle assessment (LCA) for each case as demonstrated in the current investigation. A LCA must consider all the associated environmental aspects and impacts of any construction process and material used in a structure (ISO 2006). Midpoint (Mp) indicators (such as tons of equivalent carbon dioxide, CO2e, and intermediate effects representing quantity of pollutants) and Endpoint (Ep) indicators (based on damage models which are more understandable but also more complex) are typically used in LCA. There is scientific agreement with regard to the environmental impacts (obtained from LCA) in vectorized profiles. It is possible to use methods that include socio-political preference factors, internally weigh the impacts, and give a single-final score from LCA. The economic pillar mainly comes from project budgets, including materials and process costs, and can also include maintenance and dismantling costs. Social/functional indicators (e.g., resistance to fire, safety against climatic agents or even aesthetic considerations) may be difficult to choose or define. A typically good solution to quantify these indicators is to perform opinion surveys on professionals in the same knowledge field and then process the results and weigh the solutions through hierarchy processes such as AHP techniques (Saaty 2008), among other procedures. All these criteria and indicators can be evaluated deterministically or probabilistically using the MIVES tool. 2.2

Case Studies and Design Criteria

The case studies selected in the current investigation are taken from Damians et al. (2017, 2018) representing four different earth retaining wall structure types ranging from conventional to modern in the order of gravity, cantilever, and mechanically stabilized earth (MSE) walls (Fig. 1). These solutions vary widely with respect to material quantities and type, construction method, structural behavior, and performance. Each solution includes different design calculations. Although not considered in this investigation, there are hybrid design solutions available as well. The gravity wall solution is a concrete structure with geometry selected to satisfy external stability modes of failure. The cantilever wall represents a (steel rebar) reinforced concrete structure with much less concrete material (but higher quality). The MSE wall alternatives use reinforced soil behind a thin concrete panel facing. In one case the soil reinforcement is polymeric geogrid and in the other it is steel grid. The geogrid strength and steel grid bar diameter vary with layer depth below the wall crest in accordance with current accepted design procedures which are based on variants of active wedge

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theory. A functional unit (FU) was defined for each one meter-width of wall face and assuming a wall height of 5 m, horizontal back slope and design life of 100 years. The soil was considered to be dry, granular (non-cohesive), with a unit weight of 20 kN/m3, friction angle of 30º for the retained backfill and foundation, and 34º for the reinforced backfill in MSE wall cases. The four proposed solutions were designed to minimize construction material quantities while achieving minimum acceptable margins of safety against internal and external modes of failure as recommended in AASHTO (2016) and ACI (2002) design codes.

Fig. 1. Geometry for (a) gravity, (b) cantilever and (c) MSE walls with steel or polymeric soil reinforcement. All dimensions in metres.

The geometrical dimensions and structural details for all wall solutions are presented in Fig. 1. The required materials, time and transportation inventory for the case

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study (H = 5 m) are shown in Table 1. The system boundaries and material quantities in each alternative have been framed by the system unit definition and determined by each design criteria and also taking into account all aspects related to material transport and construction works. Detailed information for other wall heights can be found in Damians et al. (2017, 2018). Table 1. Materials, time and transportation inventory for H = 5 m wall case. Category

Related items and components

Units Wall type Gravity Cantilever MSE wall: steel & polymeric 3 6.7 2.3 0.05 Structural Concrete m materials Rebar tones N/A 0.27 N/A Precast panels m2 N/A N/A 5.6 Soil reinforcement Kg N/A N/A 54 (steel grid) 10 (geogrid) Soil materials Backfill (reinforced and/or tones 23.3 35.9 61.8 retained) Earthworks Excavation m3 1.2 0.9 3.9 Backfilling and h 1500 2800 9200 compaction time Transportation Concrete (and steel) km 10 10 N/A Panels and reinforcement km N/A N/A 100 Selected backfill km N/A N/A 10

3 Sustainablity Model for Earth Retaining Walls 3.1

General Requirements (Pillars), Stakeholder Scenarios and Value Functions

The simplified decision-making tree flow-chart generated in this investigation is presented in Fig. 2 (Damians et al. 2018). Different weighting multiplier combinations were applied to each sustainability pillar according to different hypothetical stakeholder group scenarios (see Table 2). Single criteria and indicator levels were also considered for each requirement category or pillar. As shown in Fig. 2, the value functions allow the transformation from indicator units to common dimensionless value units from 0 to 1. A value function having the exponential decay formulation is available in the MIVES software toolkit, allowing different shapes to be captured by selection of different parameters (Alarcón et al. 2011). In the current study, sensitivity analyses were carried out assuming three different shapes of the value functions to transform the indicators for the environmental and economic pillars: convex, S-shape and linear (see Fig. 3). The non-linear functions (convex and S-shape types) favor best solutions and discriminate against relatively poor solutions. For the S-shape function, indicator entry units close to the “best”

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Fig. 2. Sustainability assessment flow chart (Damians et al. 2018). Table 2. Requirement weighting scenarios considered. Stakeholder group/pillar weighting scenario A B C D

Weightings for Wrequirement Environmental 1/3 2/3 1/6 1/6

requirement levels (pillars): Economic 1/3 1/6 2/3 1/6

Societal and functional 1/3 1/6 1/6 2/3

alternative returns the most favorable value (i.e., Vindicator = 1). This is a way to identify the group of best alternatives, not just a single option. Table 3 presents the summary of all cases analyzed in this study according to the combinations of value function types considered for each pillar of sustainability. 3.2

Environmental Requirement

SimaPro software (Pré Consultants bv 2010) was used to perform the LCA. The Ecoinvent v3.1 database and ReCiPe (ReCiPe 2014) method in the SimaPro software package was used to compute final scores (i.e., simplified Midpoint and Endpoint indicators). Ep indicators, related to the expected impact on human health, eco-systems, and natural resources were selected. Complete analysis and full details of the case studies are provided in Damians et al. (2017). A decreasing convex/S-shape/linear value function for the Ep final score was used over the [1, 0] range; thus high LCA Endpoint values in Fig. 3a translate to low value function multipliers (i.e., the higher the value, the less environmentally friendly the solution option). In the current study, two additional constraints on value function outcomes were imposed: Endpoint values

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Fig. 3. Value functions assumed and limits for (a) environmental and (b) economic level analyses.

that were twice the average value of all four solution alternatives for the given wall height case (H = 5 m) were assigned a score of zero, and endpoint values with a score less than 50% of the average of all solutions were assigned a value of 1. This strategy was used to ensure that obvious low environmental-impact cases were assigned a maximum value and solutions with very high environmental impact with respect to the other solutions were not rewarded with a non-zero value. As noted earlier, both indicator- and criterion-level weighting factors were assigned a value of 1 in the current study.

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Table 3. Summary of cases analyzed according to the combinations of value function types considered. C: Convex, S: S-shape, L: Linear (see Fig. 3). Case: Nomenclature Value function Environmental 1: CCL Convex 2: CSL 3: CLL 4: SCL S-Shape 5: SSL 6: SLL 7: LCL Linear 8: LSL 9: LLL

3.3

type for requirement levels (pillars) Economic Societal and functional Convex Linear S-Shape Linear Convex S-Shape Linear Convex S-Shape Linear

Economic Requirement

With regard to the major material quantities for each wall type in Fig. 1, an economic wall construction inventory was developed. ITEC-BEDEC (ITEC 2014) was the reference database adopted together with representative project material costs. Some construction costs were adjusted to better match actual market conditions. For example, for most of the components, the economic database already includes typical transportation and installation costs, and not only manufacturing costs. In MSE wall cases, facing panels and selected backfill transportation distances were taken to be 100 km and 10 km, respectively. Variability of common material quantities and costing processes were considered by using a representative range of costs for the construction of each alternative (detailed explanation can be found in Damians et al. 2018). After applying these cost variation assumptions for all material components and related processes, the density function of risk quantification in terms of a triangular frequency distribution function for each wall was obtained and then used to generate random cost values for a wall solution using Monte Carlo simulation. Each random cost variable was converted to an indicator value by using the economic requirement value function (transformation) explained previously (see Fig. 3b). The result is an array of random Vindicator values between 1 and 0. With this probabilistic scoring process from the economic budgets, high and low boundary values can be obtained with respect to the mode-price value, and a probabilistic average value deduced to obtain the final economic score of each alternative studied. The mode value of this array of random values was taken as Vindicator for the economic level in the sustainability assessment flow chart in Fig. 2. Similar types of value functions as the ones used in the environmental level analyses were used for the economic value function and for the same reasons (Fig. 3b) (i.e., reward low-cost solutions and discourage high-cost solutions). A value of 1 was assigned to the minimum possible cost of all solution options. Maximum cost outcomes that were more than 50% of the mode value of the cheapest solution were assigned a

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value of 1. Since criteria and requirement weighting factors were taken as 1 in this study, then Vfinal = Vrequirement = Vindicator. 3.4

Societal/Functional Requirement

To quantify the social impact requirement, it was important to identify the main civil engineering concerns with respect to the candidate earth retaining wall alternatives. To accomplish this, an opinion survey was conducted and compiled from 200 undergraduate Civil/Construction Engineering students and 50 graduate students enrolled in Geotechnical Engineering design classes at the Universitat Politècnica de Catalunya (UPC-BarcelonaTech). The survey results were used to identify and weight the general concerns for each retaining wall structure option. Relevant criteria and indicators were thus obtained. All the indicators were filtered and then evaluated by specialized professionals to ensure that all alternative wall solutions satisfy the same functional requirements. The average of the scores for each of the three survey populations was computed, and then a final Vrequirement score was assigned to each wall type by weighting the score for each group according to 20% for undergraduate students, 30% for graduate students and 50% for the experts. This procedure enabled a direct weighting of indicators and criteria that was judged sufficient for this simplified example. The value function that appeared in this string of calculations was linear for all cases in this study, with a score of 1 transformed to 0 and a score of 5 transformed to 1. Weighting values for indicator and criteria items were deduced from survey results described in the next section. The calculation of Vrequirement for the societal/function level follows the flow path shown in Fig. 2. The value function that appeared in this string of calculations was fixed as linear (L) in all cases analyzed in the current study, with a score of 1 transformed to 0 and a score of 5 transformed to 1. Weighting values for indicator and criteria items were deduced from survey results and are described next.

4 Sustainability Assessment Results Figure 4 shows the environmental Midpoint (Mp) indicators for the global warning potential (tones of CO2e) and cumulative energy demand (GJ) obtained for the defined functional unit for the case studies (i.e., 5 m-high earth retaining wall structures). The results obtained from the single-final score using Endpoint (Ep) model indicators analysis are identified in Fig. 5. The amount of material and the different environmental %-effect of each component (  3%) are shown in each component breakdown for each wall alternative. The results show that the MSE wall options resulted in better Mp and Ep values, with similar trends/differences obtained between alternatives for H = 5 mhigh structures. Additional results from the LCA have been reported by Damians et al. (2017). The results for the economic pillar (requirement) analyses are presented in Fig. 6. Construction work sequence and construction work type have been selected and are included in the results. Figure 6 also includes the computed cost ranges for each alternative according to the total minimum, maximum and mode cost values. The

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4

Global warming potential (tones CO2 eq)

a) GWP

-Reinforcing steel (Cantilever) -Soil reinforcement (MSEW): steel or polymeric

3 Construction activities and materials transportation

2

1

b) CED

Cumulative energy demand (GJ)

0 25 20

Backfill soil -Concrete (Gravity & Cantilever) -Precast panels (MSEW): concrete plus steel rebar and connections

15 10 5 0

Gravity Cantilever MSEW: MSEW: Steel Polymeric Fig. 4. Midpoint single score results: (a) Global warning potential and (b) Cumulative energy demand.

differences between the maximum and minimum %-values influences the resultant value transformation from the value function shape selected through the probabilistic process. Table 4 presents the criteria, indicators and weightings used in the societal pillar analysis. Again, the MSE wall solutions are judged best, although traditional solutions have an advantage in about half of the criteria categories. The table shows that the weighting assigned to each indicator is critical to optimal solution outcomes. The final pillars and MIVES scores are presented in Figs. 7 and 8, according to the stakeholder group and related sustainability pillar weighting scenario assumed (see Table 2). It can be observed that using the proposed model and methodology, the MSE wall with polymeric-geogrid reinforcement gives the highest (best) rating for

Assessment of Earth Retaining Wall Sustainability

Endpoint - Single score (Pt)

250 200

Natural resources Ecosystems

23%

Human health

150

26%

21%

100 50

34%

51% 45%

35% 22% 43%

33% 23% 44%

0

Gravity Cantilever MSEW: MSEW: Steel Polymeric Fig. 5. Endpoint single score results and contributions.

Cost per metre running length of wall (€)

La: Labour; Ce: Construction equipment; Mt: Materials; Tr: Transport

1200

Tr

Tr

1000 800

Mt

Tr

Mt Mt

600 Ce 400 200

Tr

Mt

Ce La

La

Ce

Ce

La

La

0

Cantilever

MSEW: MSEW: Steel Polymeric

Gravity Costs range (%): (-12;+18%) (-12;+18%) (-14;+24%) (-14;+23%) Fig. 6. Economic inventory using mode prices.

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Table 4. Societal and functional criteria and indicators: survey form results (Damians et al. 2018). Criteria (& weightings)

Indicatorsa (& weightings)

Wall type Gravity Cantilever

MSE MSE wall: wall: Polymeric Steel Marketing 1 (25%) **** ***** *** ** considerations (25%) 2 (15%) ***** **** *** *** 3 (10%) * * **** ***** 4 (25%) ***** ***** ** *** 5 (15%) * * **** *** 6 (10%) **** **** *** *** Design and construction 7 (20%) ***** *** *** ** methodology (25%) 8 (25%) *** *** *** *** 9 (30%) ** ** **** ***** 10 (25%) ** ** *** **** Aesthetics (15%) 11 (100%) ** *** **** **** Reliability (20%) 12 (25%) ** ** *** *** 13 (25%) ***** ***** *** *** 14 (25%) ***** **** *** *** 15 (25%) ** ** *** *** Resilience (15%) 16 (50%) ** * ** ** 17 (50%) ** ** *** **** a Indicator numbers: 1-Acceptance of wall type; 2-Labour requirements (less is better); 3Research and development required (more is better); 4-Use of local materials and technology (more is better); 5-Specialist/trained workers required (more is better); 6-Land use (less is better); 7-Ease of design; 8-Safety during construction; 9-Ease of construction; 10-Duration of construction; 11-Aesthetics; 12-Ease of repair (more is better); 13-Ease of routine maintenance (more is better); 14-Expectation of satisfactory performance; 15-Consequences of poor performance requiring repair (low is better); 16-Flexibility to design changes during construction; 17-Robustness against site conditions changes from design specifications (e.g., water).

sustainability assessment in almost all stakeholder group scenarios and value function decay shapes considered in this investigation. If environmental issues are the most important concern of stakeholders, then the MSE solutions are the best solution by a substantial margin (see Fig. 7b). However, for pillar weighting scenario C (i.e., economic stakeholder group) conventional wall alternatives become viable and cantilever walls are the best alternative if the S-shape value function is used to convert costs to indicator values (see Fig. 8a).

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Fig. 7. Sustainability results for different weighting scenarios.(a) A: 1/3-En. 1/3-Ec. 1/3-So. and (b) B: 2/3-En. 1/6-Ec. 1/6-So.

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Fig. 8. Sustainability results for different weighting scenarios. (a) C: 1/6-En. 2/3-Ec. 1/6-So and (b) D: 1/6-En. 1/6-Ec. 2/3-So.

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5 Conclusions Sustainability assessment is becoming a key requirement for selection of the best solution in civil engineering works. This paper presents a simplified but promising approach for sustainability assessment. The model is based on the MIVES methodology using value theory and multi-attribute assumptions. The general approach is applied to four different 5 m-high earth retaining walls that perform the same function. The analysis results show that MSE wall solutions are most often the best option in each sustainability pillar category (environmental, economic and societal/functional) compared to conventional gravity and cantilever wall solutions, and thus most often the best final choice when scores from each pillar were aggregated to a final score. Nevertheless, different scenarios are presented with regard to stakeholder preferences for the relative importance of the three sustainability pillars. When cost is weighted most highly of the three pillars, then the conventional wall solutions give best MIVES score for walls of 5-m height. In this scenario, the cantilever wall alternative gives the highest (best) score if the S-shape value function is selected to assess the economic pillar. If environmental issues are the most important concern of stakeholders, then the MSE wall solutions are the best solution by a substantial margin. The MSE wall solutions were shown to be the best type of structure based on the MIVES methodology mainly because a large volume of the structure is comprised of soil rather than concrete and reinforcement steel.

References Aguado, A., del Caño, A., de la Cruz, M.P., Gomez, D., Josa, A.: Sustainability assessment of concrete structures within the Spanish structural concrete code. ASCE J. Constr. Eng. Manag. 138(2), 268–276 (2012) Alarcón, B., Aguado, A., Manga, R., Josa, A.: A value function for assessing sustainability: application to industrial buildings. Sustainability 3, 35–50 (2011) AASHTO: AASHTO LRFD Bridge Design Specifications, 6th edn. American Association of State Highway and Transportation Officials, Washington, DC (2016) ACI Committee 318: Building Code Requirements for Structural Concrete (ACI 318-02) and Commentary (318R-02) American Concrete Institute, Farmington, MI (2002) Damians, I.P., Bathurst, R.J., Adroguer, E., Josa, A., Lloret, A.: Environmental assessment of earth retaining wall structures. ICE Environ. Geotech. (2017). https://doi.org/10.1680/jenge. 15.00040 Damians, I.P., Bathurst, R.J., Adroguer, E., Josa, A., Lloret, A.: Sustainability assessment of earth-retaining wall structures. Environ. Geotech. (2018). https://doi.org/10.1680/jenge.16. 00004 ISO: Environmental management, Life Cycle Assessment, Principles and Framework (ISO 14000:2006). International Organization for Standardization, Geneva (2006) ITEC: Banco estructurado de datos de elementos constructivos (BEDEC). Foundation Catalonia Institute of Construction Technology, Barcelona (2014). http://www.itec.es/home/index.asp Jones, C.J.F.P.: Guide to Reinforced Fill Structure and Slope Design (Geoguide 6). Geotechnical Engineering Office, Civil Engineering Department, Hong Kong (2002)

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Josa, A., San José, T., Cuadrado, J.: El caso de la EHE. Jornada sobre Sostenibilidad en la tecnología del hormigón: MI-VES, una herramienta de apoyo a la toma de decisiones, pp. 84– 95. Department of Construction Engineering, UPC-BarcelonaTech, Barcelona (2008) Josa, A., Alavedra, P.: El concepto de sostenibilidad. In: Losada, R., Rojí, E., Cuadrado, J. (eds.) La medida de la sostenibilidad en edificación industrial: MIVES, pp. 59–70. UPV, UPC, Labein-Tecnalia, Bilbao (2006) Pré Consultants bv. SimaPro software, v.8. Amersfoort (2010) ReCiPe: ReCiPe (2014). http://www.lcia-recipe.net/ Saaty, T.: Decision making with the analytic hierarchy process. Int. J. Serv. Sci. 1, 83–98 (2008) WCED: Our common future, Report of the World Commission on Environment and Development, The Brundtland Commission, United Nations Documents, NGO Committee on Education, Oxford University Press, Oxford (1987)

Numerical Study of Passive Earth Pressure on Retaining Walls Meriem Fakhreddine Bouali1 and Mounir Bouassida2(&) 1

Faculty of Sciences and Technology, Laboratory of Management, Maintenance and Rehabilitation of Facilities and Urban Infrastructures, University of Souk Ahras, Souk Ahras, Algeria [email protected] 2 Université de Tunis El Manar- Ecole Nationale d’Ingénieurs de Tunis, LR14ES03-Ingénierie Géotechnique et Géorisque, BP 37 Le Belvédère, 1002 Tunis, Tunisia [email protected]

Abstract. The determination of passive earth pressure by a vertical rigid wall on a horizontal backfill made up of cohesion less material is studied. Two types of movement were considered to generate the limit equilibrium of passive earth: translation movement, and rotation around the top of the rigid wall. Analysis consisted of series of 2D finite difference FLAC code. Parametric study included the effect of type of wall movement, soil-wall interface friction angle for analyzing the distribution of passive pressure and the location of resulting passive. Predicted results were found in good agreement with measurements obtained from scaled test models and full-scale retaining structure. When the translation movement is assumed, the distribution of passive earth pressure is overall linear with depth. In turn, when the mode of rotation around the top is considered, the variation of passive pressure with depth is rather non-linear. Further, the respective resultant force of passive pressure exerted on the retaining wall are located at different depth which are both lesser than one-third of the height of the wall. The location of passive force is also depending on the soil-wall interface friction angle. Keywords: Backfill
Cohesionless
Passive pressure
Rigid wall
Rotation
Translation

1 Introduction For studying limit equilibrium cases, i.e. passive and active states, two theories had been formulated quite earlier by Coulomb (1776) and Rankine (1857). Coulomb’s (1776) theory considered plane failure in the retained soil which is in frictional contact with a rigid wall. The active, or passive, force is determined after the wedge soil limit equilibrium method. Classical Rankine’s theory (1857) was formulated to determine passive and active earth pressures on a vertical rigid curtain subjected to translation movement by assuming smooth contact between the soil and retaining structure. Lancellota (2002) provided an analytical solution for the passive earth pressure coefficients based on the lower bound theorem of plasticity. Reddy et al. (2013) © Springer Nature Switzerland AG 2020 P. Pinto et al. (Eds.): GeoMEast 2019, SUCI, pp. 89–105, 2020. https://doi.org/10.1007/978-3-030-34190-9_8

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implemented numerically the limit equilibrium method for determining the passive earth pressure. From the horizontal stress distribution, the passive or active earth force is calculated. Due to linear variation with depth of horizontal stress, this force is applied at depth equals two-third of the height of retaining structure. Theoretical and experimental studies showed that smooth contact assumption between the soil and retaining structure is rather unrealistic when the wall facing is rough. Terzaghi (1941); Terzaghi (1936). Terzaghi et al. (1996) reported that, in case of passive earth state, smooth contact assumption often leads to non-conservative estimations of earth pressures. The effect of existing roughness between a frictional material and retaining structure, especially when earth pressure is mobilized, cannot be neglected. Earth pressures exerted by, or against, a vertical retaining wall is often depending on wall, or soil, displacement. In case of rigid walls, the type of displacement revealed high influence parameter for estimating earth pressures. Bang (1985) proposed an analytical estimation of the magnitude and distribution of lateral earth pressure exerted by cohesionless soil retained by a wall undergoing rotation about its base by which both the type of wall movement and magnitude of wall displacement are taken into account. In particular, Bang (1985) stated that the distribution of passive soil pressure on rigid retaining walls varies with the magnitude of horizontal displacement and type of wall displacement, i.e. translation and/or rotation around the top or the basis. Coefficients of passive earth pressure Kp, and the normalized depth of point of application (centroïd), h/H, represent two key parameters for the calculation of passive earth pressures. The assumed location of resultant of force at one-third of the wall’s height still holds for the design. The inherent assumption of linear variation of earth pressure with depth, for all types of retaining structures, is yet questionable. For this purpose, an insight should be made in regard to the determination of passive earth pressure by analyzing the mode of wall movement. Based on numerous existing contributions, it is well agreed that the magnitude, the direction and the location of the force of passive earth pressure are depending on the mode of movement, Peng et al. (2012). Aided by the development of computer tools and variety of numerical methods, many researchers investigated numerical modeling to better capture the determination of earth pressure. As an example, Roscoe (1970) studied the computation of passive earth pressure of a rigid wall rotating either about its top or about its toe. From this early contribution, it had been demonstrated that the magnitude; the direction and the location of the applied resultant force of passive pressure are generally depending on the type of wall movement. The effect of wall movement was always adopted for analyzing the stability of retaining walls. Since then, and following Clough and Duncan (1971) contribution, further developments were published, especially, to draw the attention concerning the mobilization of friction component on the interface between backfill material and retaining structure for each increment of displacement, Matsuzawa and Hazarika, (1996). This paper discusses the variation of the distribution of passive earth pressure against a rigid retaining wall subjected to a translation and rotation around the top. The resultant passive force and its point of application are also determined by the explicit

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finite difference code FLAC (Fast Langrangian Analyses of Continua, 2005). The usefulness of proposed results for the translation mode is highlighted, and then the influence of a rotation by top are presented, commented and, then, compared with existing results.

2 Problem Statement – Numerical Modelling Consider a rigid vertical wall, of height H = 1 m, which mobilizes the passive earth pressures in a purely frictional material, with free of stress horizontal surface. Plane strain numerical calculations are implemented to evaluate the coefficient of passive earth pressure and the location of the corresponding total passive force for two cases. The first case considers a vertical rigid wall subjected to translation movement (T) shown in Fig. 1a (S denotes the assumed uniform wall’s horizontal displacement). In the second case, a rotation around the top (RT) is investigated as shown in Fig. 1b, where h is the angle of rotation of rigid wall. For the two cases, the wall displacements mobilize the limit state traduced by the passive earth pressure in the purely frictional material. S

H=1m

H=1m θ

(a) TranslaƟon

(b) RotaƟon around top Fig. 1. Studied movements of retaining rigid wall

A parametric study is conducted to analyze the influences of material-wall, interface, friction angle d, friction angle of backfill material u and the type of wall movement, i.e. translation and rotation around the top is analyzed. The two-dimensional finite difference element discretization using FLAC (2005) is used to model the backfill material in plane strain condition. Backfill material is assumed dry, homogenous and cohesion-less. The self-weight of retained material is, then, the unique loading parameter. Preliminary simulations are conducted to determine the suitable geometry of retaining wall model by testing the influence of the adopted mesh and boundary conditions. The adopted numerical model is shown in Fig. 2 with following boundary conditions: Along the right-side vertical border (x = 8H; 0 < y < 6H): only vertical displacement is allowed.

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Fig. 2. Geometry and boundary conditions of numerical model

Along the left-side vertical border (x = 0; 0 < y < 6H): two different conditions are considered: From ground surface to depth equals to the rigid wall height H (x = 0; 0 < y < H), imposed uniform horizontal displacement is applied in case of translation movement (Fig. 1a); imposed orthoradial displacement is applied in case of rotation around the top movement. From the wall base to the bottom of numerical modeling (x = 0; H < y < 6H): only vertical displacement is allowed. At the bottom of numerical modeling (y = 0; 0 < x < 8H): horizontal border horizontal and vertical displacements (Fig. 2) are prevented. Several numerical tests were compiled to choose the suitable dimensions of the numerical model, they were chosen such that the developed failure surfaces within the backfill material do not cross the vertical right-hand side border. Benmeddour et al. (2012) reported that the suitable dimensions of the model in the horizontal (x) and vertical (y) directions are eight and six times the wall’s height, respectively. Computations by implementing larger dimensions of numerical model indicated that predicted passive or active earth pressure are not influenced when extending the boundaries far away from the rigid wall. The adopted mesh geometry in the present analysis is eight and six times the wall height in horizontal (x) and vertical (y) directions, respectively. A refined mesh neighboring the soil-wall interface to see much better significant gradients’ displacement at left top corner of model shown in Fig. 2 (0 < x < 2.25H and 3.75H < y < 6H). To simulate properly the experimental condition, interface elements should be introduced in the numerical study. Hence, the frictional contact at the interface between the wall and backfill material is modelled by using joints with zero thickness (Matsuzawa and Hazarika, 1996). Interface elements of the backfill material-rigid wall

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contact are characterized by Coulomb shear-strength criterion. The cohesionless interface between the backfill material and rigid wall is characterized by the friction angle d which takes the values from zero, to simulate the case of smooth contact, and values d = (1/3; 1/2; 2/3; 1)u is to describe the roughness of soil-wall contact. This interface is also characterized by a normal stiffness kn = 103 MPa/m, and a shear stiffness ks = 103 MPa/m. It was adopted that kn and ks be set to ten times the equivalent stiffness of the stiffest neighboring element (FLAC, 2005). The linearly elastic-perfectly plastic constitutive model, obeying to Mohr-Coulomb criterion with associative flow rule, is adopted. As reported by De Brost (1984) and Benmeddour et al. (2012); analyses adopting associative flow rule (w = u) show a ‘smooth’ and stable load-displacement response. Whilst when considering analyses with non-associative flow rule (w = 0), unstable displacement gradients are observed. Benmeddour et al., (2012) also reported the insignificant influence of dilatancy angle, w, on Kp value when u < 30°. Adopted elastic moduli are: shear modulus G = 10 MPa, and bulk modulus K = 30 MPa, which correspond to Young’s modulus E = 27 MPa and a Poisson’s ratio m = 0.35, unit weight c = 20 kN/m3 (Benmeddour et al., 2012). Note that the values of the elastic parameters have negligible effect on the magnitude of passive earth pressure coefficient which does not depend on the elastic and soil parameters (Salençon, 1990). The friction angle of backfill material is varied from 20° to 40° by 5° increments. In order to generate the passive failure, the rigid retaining wall moves towards the homogeneous cohesion less soil by imposing a horizontal displacement along the wall height. Two modes of movement are considered; a translation (T mode) and a rotation around the top (RT mode). The numerical model is simulated by imposing a very low velocity (10−6 m/step) at all nodes of the rigid wall in the case of translation (T) movement (Fig. 3a) and an angle of rotation around the top (h = 2.10−4 rad) for the rotation movement (Fig. 3b). It is noted that meshed zones in Figs. 3a and 3b correspond to the area located at left top corner of Fig. 2 with of depth 1.65H (vertical direction) and 3.25H (horizontal direction). Therefore, the induced failure mechanisms are not affected by the actual borders of adopted numerical model shown in Fig. 2. The corresponding horizontal movement of the rigid wall induces the progressive increase in soil passive pressure up to the occurrence of steady plastic state flow within the soil. This latter corresponds to the failure phase for which the resultant of passive earth pressures stabilizes at some value, considered as the ultimate passive load (Fig. 4). At this moment, the magnitude and the centroid of the resultant of passive earth pressures exerted on the wall are recorded. The asymptotic value of the ultimate force corresponds to the normal resultant of passive earth force (Ppx). Hence, the passive earth pressure coefficient Kp is calculated from Eq. (1): Kp = 2Ppx /c
H2
cos d

ð1Þ

Ppx: denotes the horizontal component of passive earth force. The centroid is the normalized depth which is equal to “h/H”; h is depth between the location of passive force and the wall base.

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FLAC (Version 7.00) LEGEND 6.200

10-Sep-18 18:49 step 26452 -1.629E-01 0.85 are susceptible to liquefy and loose soils with 12 < IP < 20 and wa/wL > 0.85 have higher strength to liquefaction and soils with IP > 20 are not liquefiable. It is important to refer that Eurocode 8 (1998b)-Part 5 considers no risk of liquefaction when the ground acceleration is less than 0.15 g in addition with one of the following conditions: (i) sands with a clay content higher than 20% and a plasticity index > 10; (ii) sands with silt content higher than 10% and N1(60) > 20; and (iii) clean sands with N1(60) > 25.

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Fig. 34. Correlation between volumetric strain and SPT (after Tokimatsu and Seed 1987)

9.1

Settlements Assessment

The susceptibility of foundations soils to densification and to excessive settlements is referred in EC8, but the assessment of expected liquefaction - induced deformation deserves more consideration. By combination of cyclic shear stress ratio and normalized SPT N-values Tokimatsu and Seed (1987) have proposed relationships with shear strain (Fig. 34). To assess the settlement of the ground due to the liquefaction of sand deposits based on the knowledge of the safety factor against liquefaction and the relative density converted to the value of N1 a chart (Fig. 35) was proposed by Ishihara (1993). 9.2

Remedial Measures

Following EC8 ground improvement against liquefaction should compact the soil or use drainage to reduce the pore water pressure. The use of pile foundations should be considered with caution due to the large forces induced in the piles. The remedial measures against liquefaction can be classified in two categories (TC4 ISSMGE 2001; INA 2001): (i) the prevention of liquefaction; and (ii) the reduction of damage to facilities due to liquefaction. The measures to prevent of occurrence of liquefaction include the improvement of soil properties or improvement of conditions for stress, deformation and pore water pressure. In practice a combination of these two methods is adopted.

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Fig. 35. Post cyclic liquefaction volumetric strain curves using CPT and SPT results (after Ishihara 1993)

The measures to reduce liquefaction induced damage to facilities include (1) to maintain stability by reinforcing structure: reinforcement of pile foundation and reinforcement of soil deformation with sheet pile and underground wall; (2) to relieve external force by softening or modifying structure: adjusting of bulk unit weight, anchorage of buried structures, flattering embankments. 9.3

Liquefaction Evaluation

The liquefaction potential evaluation was performed only by field tests taking into account the disturbance that occurs during sampling of sandy materials. In this analysis attention was drawn for SPT and CPT tests as the seismic tests have only been used when soil contains gravel particles. The shear values were computed from a total stresses model, that gave results on the conservative side using the code “SHAKE 2000”. Just as an example Fig. 36 illustrates the differences between the total stress model and an analysis in effective stresses using the computer program DYNAFLOW for the Vasco da Gama bridge in Tagus river and with the same type of alluvia materials. Corrections related with SPT test results due to the depth effect and the equipment were performed following the recommendations of EC8 (1998b). The sieve curves of materials a1 and a2. are shown in Figs. 37 and 38.

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Fig. 36. Equivalent shear stresses computed from SHAKE and DYNAFLOW codes (after Seco e Pinto and Oliveira 1998)

Fig. 37. Sieve curves for material a1

Fig. 38. Sieve curves for material a2

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Taking into account that we are dealing with underwater materials, the sieve curves exhibit percentages of fines lower than in reality, as a consequence of the washing effect during the sampling(Seco e Pinto et al. 1997). The liquefaction potential evaluation was given in tables and the columns have included the following data: (i) columns 1 to 4, reference to the pier, type of test (SPT or CPT), depth of the test and thickness of the layer; (ii) columns 5 and 6, values of Nm (SPT) and (qc)m (CPT); (iii) columns 7 and 8, effective overburden pressure (r’o) and correction factor (CN); (iv) columns 9 and 10, normalised values N1 (60) (SPT) (for shallow soils due to disturbance effects reduced CN values were considered) and (qc)1 (CPT); (v) column 11, sequiv. (equivalent shear stress value computed for action type 2 related with the highest magnitude 7.5); (vi) column 12 (s/r’o ratio value), column 13 (s/r’o ratio value with a safety factor of 1.1), column 14 (s/r’o ratio value with the safety factor of 1.25); (vii) column 15, Ref. (reference of the analysed SPT or CPT value); (viii) column 16, liquefaction susceptibility analysis. Taking into account the dilatant behavior of the material observed in the CPT tests and the values of the pore pressures developed in the cyclic torsional shear tests, where the registered values of the pore pressures rarely reach the value of 80%, being frequently below 60%, a safety factor of 1.1 can be considered sufficient. Nevertheless, at the present case, a conservative analysis was performed, with a safety factor of 1.25 being adopted, as recommended in EC8, Part 8. 5 (1998b) (Table 10). Table 11 presents an application of liquefaction evaluation for material a1 and material a2. The liquefaction potential evaluation, by SPT and CPT tests, is shown in Figs. 39 and 40. Taking into account the Figs. 34 and 35 the estimated settlements of materials a1 and a2 are between 40 mm to 150 mm.

10 Pile Load Tests 10.1

Introduction

Following Eurocode 7 (1997) pile design can be performed by: – – – –

prescriptives measures and comparable experience; design models; use of experimental models and load tests; observational method. The piles of Leziria bridge were designed by:

(i) design models; (ii) pile load tests that have given information about the characteristics of gravel materials and techniques of driving the metallic casings; (iii) comparable experience. Pile load tests were performed with the following purposes: (i) to determine the response of a representative pile and the surrounding ground to load, both in terms of settlements and limit load;

Table 10. Evaluation of liquefaction potential material a1 and material a2

16.8–25.1 24.3–31.3 0.0–4.2 4.2–7.4 7.4–9.6 24.6–27.6 0.0–3.6 3.6–6.2 0.0–4.5 26.0–28.8 0–5.4 24.1–25.0 25.0–29.2

8.3 7.0 4.2 3.2 2.2 3.0 3.6 2.6 4.5 2.8 5.4 0.9 4.2

Nm - SPT value N1 (60) - Normalized SPT value (qc)m - CPT cone resistance value (qc)1 - Normalized CPT cone resistance r’0 - Effective overburden pressure sequiv. - Equivalent cyclic shear stress CN - Correction factor for overburden pressure L - Liquefaction N.L - No liquefaction

“ “ “ “ “

S1B S2B-2 S3B-1 S3B-2 S3B-3 S3B-4 S4B-1 S4B-2 S5B-1 S5B-2 S6B-1 S6B-2 S6B-3

44 23 3 6 12 26 4 3 3 31 2 5 17

– – 0.5 0.52 0.65 – 0.5 0.52 0.5 – 0.5 – – 139.1 215.4 33.9 66.4 89.4 200.2 31.2 58.5 20.3 191.1 24.3 164.2 188.1

0.8 0.7 1.0 1.2 1.1 0.7 1.0 1.0 1.0 0.7 1.0 0.8 0.7

37 16 3 7 13 18 4 3 3 22 2 4 12

– – 0.5 0.6 0.71 – 0.5 0.52 0.5 – 0.5 – – 39 55 7.6 19.2 26.3 52.0 6.6 16.5 8.3 55.1 9.7 48.7 54.4

0.29 0.26 0.22 0.29 0.29 0.26 0.21 0.28 0.41 0.29 0.40 0.30 0.29

0.36 0.32 0.28 0.36 0.37 0.32 0.26 0.35 0.51 0.36 0.50 0.37 0.36

A2 A2 A2 A1 A1 A2 A2 A2 A2 A2 A2 A2 A2

N.L L L L L L L L L N.L L L L

(1) (2) (3) (4) (5) (6) (7) (8) (9) (10) (11) (12) (13) (14) (15) (16) Pier No of bore Depth (m) Thickness (m) Nm (qc)m (MPa) r’o (kPa) CN N1 (60) (qc)1 (MPa) sequiv. (kPa) s/r’o s/r’ox 1,1 s/r’ox 1,25 Mat. Remarks hole or CPT

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(ii) to check the performance of individual piles and to allow judgment of the overall pile foundation; iii) to assess the suitability of the construction method. Load tests were carried out on trial piles which were built for test purposes before the final design. The results of load tests were used to calibrate the design parameters and so to optimize the suggested values for pile lengths, based only on the interpretation of site investigation and laboratory and in situ test results. The number of pile tests were selected taking into consideration the following aspects: – – – –

the ground condition and the spatial variation; the geotechnical category of the structure; past experience related the use of same type of piles in same ground conditions; planning of the works.

The experimental piles for static and dynamic tests were located at km 8 + 200 where the pile was embedded 1 diameter in the Miocene, at km 7 + 900 where the pile was embedded 3 diameters in the gravel materials, and at km 5 + 400 where the pile was embedded 3 diameters in the Miocene. Table 11 gives a summary o pile type and location. In each place a 800 mm diameter pile was built for static test, two reaction piles with 1500 mm of diameter, 3.5 m apart from the pile test, and a fourth 800 mm diameter pile, 5.5 m apart from the first pile, for dynamic test. To perform pile load tests 7 piles 1.5 m diameter and 7 piles 0.8 m diameter piles were built. 10.2

Vertical Pile Load Tests

The methodology to perform static vertical pile load tests has followed “Axial Pile Loading Test, Suggested Method” recommended by ISSMGE and published in “ASTM D1143 (1981). The purpose was to incorporate the contribution of all the ground layers and their influence in the deformations until a depth of 5 diameters, unless the bedrock was situated at upper level. Vertical load tests were performed on 3 piles. For the vertical load test the following equipments were installed: 2 mechanical dial gauges, electrical displacement transducers (Fig. 41) with removable extensometers (Fig. 42), with a resolution of 10−6, and anchors, 1 temperature sensor, 1 tilmeter, 1 hydraulically operated pump, 2 hydraulic jacks and 1 optical level. A general view for vertical pile load tests is presented in Fig. 43. For the vertical pile load tests a maximum load of 9100 kN was applied, i.e. 3.25 times the service load. The loads were applied in two cycles of load and unload, with a maximum load of service load for the first cycle and the loads were applied in 4 increments. In the second cycle the loads were applied in 19 increments.

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Table 11. Summary of pile type and location Piles (km) 5 + 400 7 + 900 8 + 200 4 + 750

Diameter (m) Pile embedding 0.8 3Ø (M) 0.8 3Ø (a3) 0.8 1Ø (M) 1.5 3Ø (M)

Type load test Vertical dynamic Vertical dynamic Vertical dynamic Horizontal dynamic

Fig. 39. Liquefaction potential evaluation by SPT tests

The number of load increments and the cycles of load and unload were defined with the purpose to reach some conclusions related to deformations, creep effects and ultimate load. The load - settlement curves for 3 pile tests are shown in Fig. 44. Failure loads were defined as settlement equal to 10% of the pile diameter, i.e. at 80 mm settlement. 10.3

Horizontal Pile Load Tests

The horizontal load tests were performed in two piles of 800 mm and 1500 mm of diameter located at km 5 + 400. The maximum load was 600 kN to mobilize a displacement of 8 cm and the loads were applied in steps of 75 kN (ICIST-IST 2005). For the horizontal load tests the following equipments were installed:

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Fig. 40. Liquefaction potential evaluation from CPT tests

– – – – – –

clinometers vibrating wire transducers load cells retrieval extensometers inclinometer tubes to measure horizontal displacements temperature device.

The loading program consisted of: 10 load increments from 50 kN to 500 kN. The load displacement curve measured is shown in Fig. 45. The measured rotations values versus loads are shown in Fig. 46. Figure 47 shows a comparison between the bending moments values obtained by the tests and by the analyses for different values of k = 2500 kPa, 5000 kPa, and 10000 kPa. 10.4

Dynamic Pile Tests

Dynamic pile tests were performed in 9 piles with diameters of 800 mm and 1500 mm. The piles were instrumented with: – 4 pairs of acelerometers (Fig. 48). – 4 transdutors – topographic equipment A dynamic test view is shown in Fig. 49. During the tests the height of the hammer fall was increasing from 0.2 m to 3.0 m in steps of 0.2 m.

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Fig. 41. Displacement transducers

Fig. 42. Recovery extensometers

The point resistance (Rb) and the lateral resistance (Rs) for pile E 800-2 is shown in Fig. 50. It is important to stress that the results of dynamic tests have confirmed the results of static tests pointing the higher contribution of the lateral resistance in comparison with the point resistance. 10.5

Part 4

“The important thing in science is not so much to obtain new facts as to discover new ways of thinking about them”. (Sir W. Bragg, British Scientist, 1968)

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Fig. 43. General view for vertical pile load tests (after Ferreira et al. 2008)

Fig. 44. Load settlement curves for vertical tests (after ICIST-IST 2005)

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Fig. 45. Measured load displacement curve for horizontal tests (after ICIST-IST 2005)

Fig. 46. Measured load rotations curve for horizontal tests (after ICIST-IST 2005)

11 Reception Tests for Piles The development and implementation of non destructive techniques of pile tests have experienced a great increment as the use of core sampling and load tests to control the final quality of the piles are very costly and can only be performed in a small number of piles. Anomalies that impair the integrity of a pile and that are expected to be identified by integrity tests include the presence of material of poorer quality than expected (locally and overall) and variations in the cross section of the shaft (e.g., crack, necking, and bulb).

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Fig. 47. Bending moments (after ICIST-IST 2005)

Fig. 48. Transducers and accelerometers

Also sonic diagraphy tests were performed and a continuous record through the length of the pile of the velocity of sonic waves between the source and the geophones introduced in two pipes attached to the pile reinforcement was done. The sound velocity in concrete is around 4000 m/s, but in the presence of anomalies, i.e. fissures, segregations or soil inclusions this value decreases. The quality of the results depends of the following requirements: (i) Use of metallic tubes with diameter between 35 and 60 mm; (ii) The number of tubes depends of the pile diameter:

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Fig. 49. Dynamic test (after Ferreira et al. 2008)

Fig. 50. Mobilized resistances (after ICIST-IST 2005)

diameter < 0,60 m = 2 tubes 0,60 m < diâmetro < 1,20 m = 3 tubes placed 120° apart diâmeter > 1,20 m = 4 tubes, as a minimum; (iii) The connection between the tubes should be done by joints; (iv) A good contact between the tube and the concrete; (v) At the bottom of the tubes a sealing should be placed to avoid the uplift of the sediments or concrete; (vi) The tubes should be connected to the pile reinforcement along the total length; (vii) The top level of the tubes should be 0.5 m above the pile head, as a minimum; (viii) The tubes should be placed vertical and parallel to the pile reinforcement; (ix) The pile test should be performed 3 days after the concreting, as a minimum. Figure 51 shows a pile view with 4 tubes. Taking into account that piles were 1.52 m diameter 4 tubes 90º apart were placed.

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In the experimental pile tests located at km 5 + 400, km 7 + 900, km 8 + 200 a verification of integrity tests by cross hole tests were performed. For piles 1.5 m diameter 4 tubes were placed.

12 Monitoring During Construction and Long Term 12.1

Introduction

The designer has the difficult task to perform a correct definition of loads and an adequate characterization of the materials for the project. It is necessary to compare the mental model with the prototype response in order to assess the structural behavior, and to decide in face of an anomalous behavior. Within this framework it is important to instrument the bridge with the following purposes: (i) (ii) (iii) (iv)

12.2

Validation of design criteria and calibration of mental model. Analysis of bridge behavior during its life cycle. Corrective measures for the rehabilitation of the structure. Cumulative experience that will be useful for the construction of more economic and safer bridges. Quantities to Be Measured

For the superstructure the measurement of the following quantities were proposed: (a) deck vertical displacements; (b) piers cross-sections rotations; (c) internal deck and piers deformations; (d) internal deck deformations due to time-dependent effects; (e) deck and stays temperatures; (f) air temperature, relative humidity and wind speed; (g) seismic and wind induced accelerations in the deck and piers; (h) forces in stays. Related with the infrastructure the following measurements were programmed: pile head displacements using electronic teodolytes and appropriate reflectors. 12.3

Warning Levels

Four warning levels were defined: (i) warning level 1 - no interruption of traffic; (ii) warning level 2 - limitation of traffic; (iii) warning level 3 - interruption of traffic; (iv) warning level 4 - decision concerning the traffic. For warning levels 1 to 3 the maintenance team can deal with the problem alone. For warning level 4 a specialist is necessary to take the decision. 12.4

Inspections

To complement the data given by the sensors placed in different sections of the bridge regular inspections should be performed. Four levels of inspection were proposed:

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Fig. 51. 4 tubes for crosshole tests in a 1.5 m diameter pile

(i) The reference situation corresponds to a detailed inspection of all parts of the structure (foundations, bearings and decks) and the measurement of all the sensors in order to characterize the initial state of the bridge before the opening to traffic; (ii) The daily inspections aimed an efficient visual checking of the superstructure (drainage systems, road surface, expansion joints, handrail, gantries, safety barriers, lighting etc.) to detect the need of small repairs; (iii) The annual inspections are related with the visual inspection of the foundations (measurements by sensors placed into the piles), supporting structures, bearings, expansion joints, superstructures and equipment; (iv) After the opening to traffic, the first detailed inspection will be done after two years. During the operation of the bridge the frequency is five years.

13 Conclusions The following conclusions can be outlined: For Vasco de Gama bridge (1) For the pile foundations each geotechnical design situation shall be verified that no relevant limit state is exceeded. (2) Limit states should be verified by one or a combination of the following methods: design by calculation, design by prescriptive measures, design by loads tests and experimental models and observational method (Eurocode 7 1997). (3) In other to improve the pile behaviour field tests with instrumented piles are highly recommended for design purposes. (4) The results of load tests performed in New Tagus bridge and Leziria bridge for design purposes have shown how they should be used to calibrate the design

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parameters, to check the performance of individual piles, to allow judgement of the overall pile foundation and to assess the suitability of the construction method. For Leziria bridge (5) The different geotechnical campaigns implemented during the Preliminary Study (1st phase and 2nd phase) and during the Basic Design have allowed the definition of different geological and geotechnical profiles. (6) The geotechnical characteristics were obtained after a balance between the results of the field and laboratory tests. (7) The geotechnical study in the Basic Design fulfills the requirements of Eurocode 7, Specification 1536 Bored Piles prepared by CEN - Committee TC 288 and the Procedures and Specifications for Piles. (8) The Leziria bridge is located in zone A of Portugal the highest seismic zone. (9) The piles were designed by (i) design models; (ii) pile load tests that have given information about the characteristics of gravel materials and techniques of driving the metallic casings; and (iii) comparable experience. (10) Static pile load tests both vertical and horizontal were carried out on trial piles to calibrate the design parameters and to optimize the pile lengths. Also dynamic pile tests were performed. (11) The liquefaction potential evaluation was performed only by CPT and SPT tests due to the disturbance that occurs during sampling of sandy materials. Both total and effective stress analyses were performed. (12) Non destructive techniques of pile tests were performed to assess the quality of piles. (13) The objectives of monitoring during construction and long term were presented. 13.1

Lessons for Tomorrow

Today there is a need to work in large teams exploring the huge capacity of computers to analyze the behavior of bridges. Innovative methods and new solutions require high reliable information and teams integrating different experts, namely seismologists, geologist, geophysics, geotechnical engineers and structures engineers. A joint effort between Owners, Decision-Makers, Researchers, Consultants, Professors, Contractors and General Public to face this challenge is needed. It is important to understand the concepts of vulnerability and resilience. Vulnerability is associated with two dimensions, one is the degree of loss or the potential loss and the second integrates the range of opportunities that people face in recovery. This concept received a great attention from Rousseau and Kant (1756). Resilience is a measure of the system’s capacity to absorb recover from a hazardous event. Includes the speed in which a system returns to its original state following a perturbation. The capacity and opportunity to recolate or to change are also key dimensions of disaster resilience. The purpose of assessing resilience is to understand how a disaster can disturb a social system and the factors that can disturb the recovery and to improve it. It is important that engineers educate themselves and the Public with scientific methods for evaluating risks incorporating the unpredictable human behavior and human errors in order to reduce disasters.

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From the analysis of past bridges incidents and accidents occurred during the earthquakes it can be noticed that all the lessons have not deserved total consideration, in order to avoid repeating the same mistakes. We need to enhance a global conscience and to develop a sustainable strategy of global compensation how to better serve our Society. The recognition of a better planning, early warning, that we should take for extreme events which will hit our civilization in the future. Plato (428-348 BC) in the Timaeus stressed that destructive events that happened in the past can happen again, sometimes with large time intervals between and for prevention and protection we should followed Egyptians example and preserve the knowledge through the writing. We should never forget the 7 Pillars of Engineering Wisdom: Practice, Precedents, Principles, Prudence, Perspicacity, Professionalism and Prediction. Following Thomas Mann we should enjoy the activities during the day, but only by performing those will allow us to sleep at the night. Also it is important to narrow the gap between the university education and the professional practice, but we should not forget that Theory without Practice is a Waste, but Practice without Theory is a Trap. Kant has stated that Nothing better that a good theory, but following Seneca Long is the way through the courses, but short through the example. I will add through a careful analysis of Case Histories. In dealing with these topics we should never forget the memorable lines of Hippocrates: “The art is long and life is short experience is fallacious and decision is difficult”.

Acknowledgments. Special thanks are due to GATTEL and NOVAPONTE for the permission to publish the results of New Tagus bridge. It is important to refer the contribution of Virgilio Rebelo and Vicente Rodrigues for geological studies. For Leziria Tagus River Bridge several construction companies and experts were involved, namely Construction Consortium and Design Group. The studies carried out by them are greatly acknowledged. The contribution of Virgilio Rebelo for geological studies is important to refer. The field investigations were carried by Geocontrole and LNEC, and the laboratory tests were performed by Geocontrole. Special thanks are due to TACE and particularly to Mr. Secundino Vilar and also to BRISA for the permission to publish this paper. It is important to refer the contributions by IST for the static pile tests and also for the dynamic pile tests. The reception tests for piles were carried out by Geosolve.

References Ambraseys, N.N.: Engineering seismology. Earthquake Eng. Struct. Dynam. 17, 1–105 (1998) ASTM D1143: Standard Test Method for Piles Under Static Axial Compressive Load (1981)

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Bozozuk, M.: Tolerate movements of bridge foundations. In: Proceedings of the 10th International Conference Soil Mechanics and Foundation Engineering, Stockholm, Balkema, Rotterdam, pp. 699–700 (1981) Burland, J.B., Wroth, C.P.: Settlement of buildings and associated damage. SOA review. In: Conference Settlement of Structures, Cambridge, pp. 611–654. Pentech Press, London (1974) Cetin, K.O., Seed, R.B., Kiureghion: Reliability based assessment of seismic soil liquefaction initiation. In: Ansal, A. (ed.) XV ICSMGE TC4 Satellite Conference on Lessons Learned from Recent Strong Earthquakes, pp. 327–332 (2001) EN 1997 Eurocode 7: Geotechnical design. Part 1 (1997) EN 1998 Eurocode 8: Design of structures for earthquake resistance (1998a) EN 1998 Eurocode 8: Design provisions for earthquake resistance of structures - Part 5 Foundations, Retaining Structures and Getechnical Aspects (1998b) Ferreira, S., Rebelo, V., Ribeiro, J.: Pile Foundations for Leziria Bridge, 11 CNG, Coimbra (2008). (in Portuguese) Gazetas, G., Mylonakis, G.: Seismic soil structure interaction: new evidence and emerging issues. In: Geotechnical Earthquake Engineering and Soil Dynamics, ASCE II, pp. 1119– 1174 (1998) ICIST-IST: Experimental Pile Load Tests. Report. New Tagus River Crossing at Carregado, August 2005 Imai, T.: P and S wave velocities of the ground in Japan. In: Proceedings of the 9th International Conference Soil Mechanics and Foundation Engineering, Tokyo, Japan (1977) INA (International Navigation Association): Seismic Design Guidelines for Port Structures. A. A. Balkema Publishers (2001) Institution of Civil Engineers: Piling – Model Procedures and Specifications, London (1978) Ishihara, K.: Soil Behaviour in Earthquakes Geotechnics. Clarendon Press, Oxford (1996) Ishihara, K.: Liquefaction and flow failure during earthquakes, 33rd Rankine Lecture. Geotechnique 43(3), 351–415 (1993) IST: Generation of Response and Displacements Spectra of Alluvial Soils of Carregado Site, February 2004 Meyerhof, G.G.: Discussion on Paper by Skemton, Peck and MacDonald- Settlements Analysis of Six Structures in Chicago and London. In: Proceedings of the Inst. (iv. Eng. 5, nº1) (1956) Moulton, L.K.: Tolerable Movement Criteria for Highway Bridges, report Nº FHWA-TS-85-228, 86 pp. Federal Highway Administratio, Washington, DC (1986) NCEER: Proceedings of NCCER Workshop on Evaluation of Liquefaction Resistance of Soils, Summary Report, Edited by T. Leslie Youd and I.M. Idriss, National Center for Earthquake Engineering Research, University of Buffalo, Technical Report NCEER-97-0022 (1997) Oliveira, J.B., Gomes, J.P., Sêco e Pinto, P.S., Pina, C.A.: Dynamic tests performed on large piles. 11 WCEE. Acapulco, Mexico, Portugal (1996) Oliveira, R., Sêco e Pinto, P., Rebelo, V., Rodrigues, V.: Geological and Geotechnical Studies for the Design of Vasco de Gama Bridge (1997). (in Portuguese). VI CNG Polshin, D.E., Tokar, R.A.: Maximum allowable non-uniform settlement of structures. In: Proceedings of 4th International Conference Soil Mechanics and Foundation Engineering, London, pp. 402–405. Butterworths Scientific Publications (1957) Portugal, A., Perry da Câmara, A., Virtuoso, F., Rebelo, V.: New Crossing Across Tagus River in Carregado. JPEE, Lisboa (2005). (in Portuguese) Rodrigues, L.F.: Methods for seismic site investigation in engineering geology. The importance of shear wave. Research thesis, LNEC (1979). (in Portuguese) RSA: Portuguese safety and actions code for buildings and bridges (1983). (in Portuguese) Sêco e Pinto, P.S.: Pile Foundations. Pile Design Following EUROCODE 7. Portuguese Society for Geotechnique (25th birthday), pp. II-1–II-67 (1997)

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Sêco e Pinto, P.S., Sousa Coutinho, A.G.F.: Deformation of Soils and Displacements of Structures, X ECSMFE, Florence, pp. 539–542 (1991) Sêco e Pinto, P.S., Oliveira, R.: A recent difficult foundation problem: the case of the new tagus bridge. In: 4th International Conference on Case Histories in Geotechnical Engineering, St. Louis, USA (1998) Sêco e Pinto, P.S., Correia, J., Vieira, A.: Evaluation of liquefaction potential of a site located in the South of Portugal. In: Proceedings of the Discussion Special Technical Session on Earthquake Geotechnical Engineering during 14th ICSMFE, Hamburg, pp. 113–124. Edited by Pedro S. Sêco e Pinto. Published by Balkema (1997) Seed, H.B., Harder, L.F.: SPT-based analysis of cyclic pore pressure generation and undrained residual strength. In: Proceedings of Memorial Symposium of H. B. Seed, vol. 2, pp. 351–376 (1990) Seed, H.B., Idriss, I.M.: Ground Motions and Soil Liquefaction During Earthquakes. Earthquake Engineering Research Institute, Oakland, California (1982) Skempton, A.W., MacDonald, D.H.: Allowable settlement of buildings. In: Proceedings of the Institute of Civil Engineers, Part 3, no. 5, pp. 727–768 (1956) Stokoe, K.H., Darendeli, M.B., Andrus, R.D., Brown, L.T.: Dynamic soil properties: laboratory, field and correlation studies. Theme Lecture. In: Proceedings of 2nd International Conference on Earthquake Geotechnical Engineering, Lisboa. Edited by Pedro Sêco e Pinto. Published by A. Balkema, vol. 3, pp. 811–845 (1999) TC4 (ISSMGE): Case Histories of Post-Liquefaction Remediation. Committee on Earthquake Geotechnical Engineering. Technical Committee for Earthquake Geotechnical Engineering (2001) TC4, ISSMFE: Manual for zonation of seismic geotechnical hazards, published by the Japanese Society of Soil Mechanics and Foundation Engineering, Tokyo (1993) Tokimatsu, K., Seed, H.B.: Simplified procedure for the evaluation of settlements in clean sands. Report nº UCB/EERC 84/16. University of California (1984) Tokimatsu, K., Kuwayama, S., Tamura, S.: Liquefaction potential evaluation based on Rayleigh wave investigation and its comparison with field behavior. In: Proceedings of the 2nd International Conference on Recent Advances in Geotechnical Earthquake Engineering and Soil Dynamics, St. Louis, 11–15 March, vol. I, pp. 357–364 (1991) Tokimatsu, K., Seed, H.B.: Evaluation of settlements in sands due to earthquake shaking. JGE, ASCE 113, 861–878 (1987) Youd, T L., Gilstrap, S.D.: Liquefaction and deformation of silty and fine-grained soils. General Report. In: Proceedings of the of 2nd International Conference on Earthquake Geotechnical Engineering, Lisboa, vol. 3, pp. 1013–1020. Edited by Pedro Sêco e Pinto. Published by A. Balkema (1999)

Enhancement of Raft Foundation Using Micro Pile Technique Ahmed T. Farid and Mostafa A. Yousef(&) Housing and Building National Research Center (HBRC), Dokki, Giza, Egypt [email protected], [email protected]

Abstract. Raft foundations are one of the shallow foundation techniques used for carrying high loads of the building especially on compressible soil formations with high differential and total settlements. In this paper, a case study of a high rise building having non-eccentric loads is established using raft foundation. There was compressible soil stratum starting from 8.0 m down to 12.5 m underneath the raft foundation. Due to the non-eccentric loads of the building and the compressible soil layer which caused the increase in the tilting of the building with time. According to the continuous of the building movement, enhancement method should be used to overcome this moving. Micro pile technique was designed and established at the tilting sides of the building. A new raft foundation was added at the top of the old foundation to insure connect of the executed micro piles with the old raft foundation. Measurements of the tilting or the movement of the building were continued before and during execution of the micro pile technique. Observations showed a good enhancement in the behavior of the foundation system after using the new system of micro piles as there was no more tilting in the building or the differential settlements.

1 Introduction The necessity for foundation enhancement is an important issue in the case of the increase of the buildings loads due to new usage or in the case of an inadequate foundations system. According to that, it is very important in the case of non-eccentric loads to calculate the expected differential or total settlements which could occur during and after the construction. The insufficient geotechnical studies lead to insufficient calculations of the expected movements during the different stages of loading which leads to multiple risks to the structures. A further need for foundation enhancements arises when the building is used in another purpose than what it was originally designed or planned. With respect to the excessive stresses under the existing foundation due to the different cases mentioned above, the original foundation is no longer adequate to transfer the additional loads to the underneath soil strata. Geotechnical engineers should find the suitable enhancement method for each case one by one after performing more tests of soil investigations to clarify the different soil characteristics and behavior. In our case of study, a high rise building located at Giza area in the Great Cairo region which an excessive stresses occurred below its foundation accompanied with higher differential and total settlements. The excessive total and differential settlement caused © Springer Nature Switzerland AG 2020 P. Pinto et al. (Eds.): GeoMEast 2019, SUCI, pp. 154–164, 2020. https://doi.org/10.1007/978-3-030-34190-9_10

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tilting to the building and hence severe cracks and/or collapse of the building can happen with time after using the building and increase of loads. The tilting of the building was increased with time as recorded and according to that; the enhancement of its raft foundation was must. Micro piles technique offers an ideal solution for enhancement of the building raft foundation which has been used widely for enhancement of foundations, underpinning, side stabilization, etc. (Plumelle 1984; Singh and Heine 1984). Their development generally is attributed to the pioneering work of Lizzi (1984). Specific analysis and design methodologies for micro-piles are limited and largely empirical (FHWA 2000; Poulos and Davis 1980). Often, empirical rules and design procedures developed for larger diameter of drilled shafts have been used for micro pile design (Bruce and Juran 1997).

2 Description of the Study Building Our study building was located at Giza district which located at the west part of Great Cairo. The building under study consists of one basement, one ground floor for the whole area of the building. Thirteen stories were established in the south-east part while eleven stories only were established in the north-west part of the building according to its architecture design. The building was constructed of reinforced concrete with a raft foundation of 140 cm in thickness which located at 3.50 m below the natural ground surface. The building area is equal to 32.0 m in length and 11.0 m in width and the building shape in plane shown in Fig. 1. The finishing of the concrete skeleton of the building of all stories was executed in nearly seven months of its start date. After that, a differential and total settlement with tilting of the building was observed. These settlements were different at each side. So, the owner started to investigate the building by a specialized consultant engineer to recover the problem. The consultant engineer advices the owner to start doing survey and geotechnical studies to investigate the reasons of settlements and tilting.

Fig. 1. Schematic diagram of building shape

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3 Geotechnical and Survey Studies The engineering study of the reasons of causing the building settlement and tilting was started directly after finishing the building skeleton. This study was performed after requested by the owner and before using the building. The study was focused into two main studies as there were no cracks observed in the structure concrete members as seen by the consultant engineer from the overall visual visit. The two main studies was to investigate what is the causes of settlements and tilting reasons by studying the soil behavior under the building and also measure the actual movements by surveying studies and to know how its value and increase with time. 3.1

Surveying Study

The surveying study was started firstly after the consultant engineer got the job of investigation as the measuring of movements of the building with every passing of any time is important. The surveying study was focusing on the tilting of the building and the differential settlements between its corners. Measurements points are fixed on the building through different locations of its plane and its height. The measurements were compared to fixed points near the building as a bench marks which is totally fixed and not affected by any load or movements. The measurements continued up to one year from the start of investigation of the building. The recorded data shows that at the start time the tilting was obvious at the east and south sides of the building where there with two floors higher than the part in the north and west part. The east tilting value was 110 mm and 180 mm for the east and south directions, respectively. After continuing measurement for one year these values reached about 270 mm in the east direction and 390 mm in the south direction. 3.2

Geotechnical Study

The geotechnical study was performed in the same time as the consultant engineer start to perform the surveying investigation to see what the causes of the tilting and movement of the building. Two boreholes up to depth of 25.0 m below the ground surface are performed in the available locations sides of the building to study the soil behavior under the raft foundation. The two boreholes give nearly the same soil stratification. The soil stratification at the site of the building consists of fill material consists of brown silty clay with sand and broken red blocks from ground surface down to depth of 2.50 m. This fill material followed by brown to gray stiff clay up to depth of 11.5 m and from 11.5 m down to 16.0 m a gray soft clay layer with a thickness of nearly 4.50 m which is the layer has a high compressibility under loads. The soft clay soil followed from 16.0 m down to the end of boring at 25.0 m with a layer of very dense medium sand. Figure 2 shows the soil stratification at the site and the characteristics of the soil formations at the site are shown in Table 1. The raft foundation of the study building was executed at depth of 3.5 m below the existing ground.

Enhancement of Raft Foundation Using Micro Pile Technique

00.0 Made ground 2.50

Stiff CLAY

11.50 Soft CLAY 16.00

Medium SAND

25.00 Fig. 2. Soil stratification at the building site Table 1. Site characteristics of the different soil stratification Characteristics

Depth (meter) 2.5–11.5 11.5–16.0 Water content % 32–34 51–54 Liquid limit % 64–65 67–69 Plastic limit % 29–31 31–32 SPT-N 10–13 2–4

16.0–25 11–14 Non Non 32–50

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4 Discussions of Causes of Tilting According to the soil tests performed on the underneath soil samples collected from the boreholes, it was found that as shown in Fig. 2, there was a compressible layer consists of soft clay having a five-meter thickness which causes higher total settlement under the foundation. In the same time, as there was a difference in heights between the northwest and the south-east sides of the building a different stresses occurred below the raft foundation which causes differential settlement beside the total settlement occurred due to the presence of the compressible layer. Soil stresses under the foundation was determined according to the actual final total loads expected after using the building to predict the total and differential settlements. The soil stresses at the south-east of the tilting side was about 280 kN/m2 while it was about 220 kN/m2 at the north-west side. The corresponding settlements at those two sides were computed according to the final actual loads to be 545 mm and 325 mm, respectively. According to that, differential settlement under the foundation occurred and increased with time due to the increase in the building stages of loading and the increase of the settlement of the underneath weak soil. This was observed and measured by the surveying study performed with time as mentioned above. Structural analysis study was carried out to ensure that the structural elements of the building are safe under the overall final load of the building or/and the measured movements. Also, as there were no cracks have been noticed in any structural elements during the investigation and that was a good situation to enhancement the support foundation of the building only in that case. The suitable method for enhancement the raft foundation and how it was executed for the building under study is discussed in the following paragraphs a deeper focus on.

5 Enhancement Method of the Raft Foundation The stresses under the raft foundation was expected to be in the different in the two sides of the buildings and ranged between 220 kN/m2 and about 280 kN/m2 for the north-west and the south-east side, respectively according the final stage of loading. These different stresses with the presence of the soft clay layer under the raft foundation causing the highly total and differential settlements under the raft foundation. The differential settlements causing the tilting of the building which can cause a severe danger with time to the building structure as the tilting was increases by 5.0 mm monthly since the finishing time of the concrete construction. Several methods were suggested to enhance the supporting foundation and reduce the differential settlements. First suggestion was to use the grouting technique to stabilize the weak soil layer by using mixture of lime and cement, but this technique is difficult to execute under the raft foundation and there was no assure for reducing the tilting and stop the movement. Second technique suggested is to use the deep foundation type to transmit the building loads to the firm layer underneath the weak soil layer. The using of the piles was difficult due to the needs of big machines with enough space for work and also consuming more time and cost which is not suitable for that building. So, the consultant engineer suggest that the good type of deep foundation is

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the micro piles as it is applicable in this case of small spaces in the basement in spite its higher cost but it can execute with a high guarantee. 5.1

Micro Pile Design

Micro piles used in the enhancement of the raft foundation were designed according to the Egyptian Code of Foundations. The bearing capacity of the micro pile depended mainly on its reinforcement and the soil characteristics. The used micro piles consist of a pile hole of 200 mm and a reinforcement steel pipe is inserted in the hole and then injected the hole with cement grouting. The initial pile load test, which is considered mandatory, controlled the final bearing capacity. However, the bond stresses between pile reinforcement and cement grout or the friction between the grout and soil should be checked. To stop the movement of the building 86 micro piles were chosen to support the raft foundations. The distribution of the micro piles is done in the sides where the tilting was observed and occurred (east and south) as shown in Fig. 3. Structural analysis was used to distribute the micro piles around the columns according to the stresses under the raft foundation (Bruce et al. 1992). The micro pile had 0.20 m in diameter and 18.0 m in depth. The reinforcement was a steel pipe 89/76 mm out/in diameter.

Fig. 3. Micro pile distribution under the raft foundation

According to Egyptian Code of Foundation (Housing & Building National Research Center 2005), pile load test was carried out to determine the allowable bearing capacity of the pile. An initial non-worked pile was installed next to the building following the same method of instillation for the underpinning piles (Wolosick 2003). Figure 4a, shows the testing load experiment of the tested pile. While Fig. 4b shows the load-settlement relationship. The test gives that the allowable micro pile load can be taken as 300 kN as the settlement at the ultimate load which twice this allowable design load was nearly about 3.0% of pile diameter.

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Fig. 4a. Micro pile- load settlement test

0

200

Load, kN 400

600

0

2 Settlement, mm

160

4

6

8

10

12 Fig. 4b. Micro pile- load- settlement relationship

800

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5.2

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Micro Pile Execution

The procedure of constructing the micro pile was achieved according to Egyptian Code of Foundations – Part 4, deep foundations. The procedure illustrates that in the start, holes should be opened through the raft foundation at the locations of micro piles illustrated above in Fig. 3. These holes could be executed by using concrete core machine or by the drilling machine of micro piles. Then, each hole should be continued into soil with the same diameter and depth. An initial grouting or cleaning the hole from the drilling fluid should be carried out in which cement grouting replaced the drilling fluid. Then the reinforcement, which consisted of the 89/76 mm steel pipe, should be inserted inside the hole. This cement grouting is considered very important to increase the end bearing capacity of the pile. The connections between micro piles and raft foundation were considered the important part of that enhancement technique for success of this method. The head of micro piles was designed to be 300 mm above the top surface of the existing raft foundation. This top part of the micro pile was prepared by steel cap to increase the connection with the raft foundation as shown in Fig. 5. A new raft with thickness 50 cm and top reinforcement was constructed above the existing raft. Therefore, the prepared head of piles could achieve the connection between micro pile and the new raft. The connection between the new and existing raft is shown in Fig. 6. Pre-design shear connectors were used between the existing and new rafts. Shear connectors between columns and the new raft were also established. Moreover, the bond stresses between the cement grout of micro piles and concrete of the existing raft play a considerable effect. High tensile steel bars were used as shear connections between the existing and the new raft foundation. These bars penetrated 250 mm through the existing raft at 40 cm center to center with 45 cm in height to reach the top reinforcement of the new raft as shown in Fig. 6.

Fig. 5. The connection between the existing, the new raft foundation and the micro piles

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Column

New Raft Foundation 50 cm in thickness Shear Connectors

Steel 10 Φ 19 / m`

R.C of Old Foundation

P.C

Micro Piles 20 cm dia

Steel Pipe 89/76

Fig. 6. The connection between the existing, the new raft foundation and the micro piles

6 Evaluation of Using Micro Piles Technique During and after the micro piles method and the new raft foundation were executed for the present building, the surveying studies was continued for these three months of execution period and one more year after that. During the three months of execution the micro piles and its new raft foundation, the movements of the building were continued. It was noticed that an additional tilting of the building was continued. The tilting of the east side was increased from 270 mm up to 285 mm before using the micro piles up to finishing them. While, the tilting was increased from 390 mm up to 410 mm in the south part of the building at the same time. This increase in the tilting measurements was regarded to the drilling of micro piles and their execution plus the increase in the new part of the raft foundation added to the structure load. After the completion of the total enhancement of raft foundation, the measuring of the building movements up to one year gives no more increase in the tilting readings which gives a good guarantee for the used technique.

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7 Conclusion The conclusion of the present study of the tilting building can be summarized in the following statements: 1. The geotechnical studies before the execution of any type of structures should be performed taking into consideration the effect of different stresses on the raft foundation which gives the movement or tilting of the building structure, especially with the presence of a compressibility soil formation underneath the foundation as seen in our case. 2. The enhancement method or technique for the raft foundation can be applicable and easy to perform before using the building and before all the total loads of the building are applied. 3. The pile load test of the micro pile which was performed to support and enhance the structural studies for calculating the micro pile load and the effective number of micro piles needed for the enhancement. 4. The micro pile technique is suitable method to use for enhancement of the raft foundation in spite of its cost because it is easy to perform and it gives more guarantee for reducing and stop the movement of the building as illustrated in our study. 5. The authors recommended that the present study for the enhancement of the raft foundation using the micro piles as mentioned above should be taken into consideration its cost compared to the height of the building and its importance. Acknowledgment. The authors would like to thank Prof. Dr. Adel Hammam (Geo-institute) and Eng. Hayel M. El-Naggar for their collaboration and assistance during performing of the present study. Special appreciated to the contractor company staff executing the micro piles for their help and patient.

References Bowles, J.E.: Foundation Analysis and Design, 5th edn. McGraw-Hill, New York (1996) Bruce, D.A., Juran, I.: Drilled and Grouted Micropiles; State of Practice Review, FHWA-RD-96016, Melean, July 1997 Bruce, D.A., Hall, C.H., Triple, R.E.: Structural underpinning by pin piles. In: Proceedings of the 17th Deep Foundation Institute Annual Meeting, New Orleans, pp. 49–78 (1992) Federal Highway Administration. Micropile Design and Construction, Report No. FHWA-NHI05-039, United States Department of Transportation, December 2005 FHWA. Micropile Design and Construction Guidelines, Implementation Manual. Publication No. FHWA-SA-97-070 (2000) Horpibulsuk, S., et al.: Underpinning for distressed building in Northest Thailand. In: 16th International Conference on Case Histories in Geotechnial Engineering, Arlington, VA, Paper No 1.02, pp. 1–9 (2008) Housing & Building National Research Center. Egyptian Code of Practice for Soil Mechanics and Foundations-Deep Foundations, vol. 4, Cairo, Egypt ( 2005) Lizzi, F.: Static Restoration of Monuments. Sage Publication, Genoa (1984)

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Plumelle, C.: Improvement of the bearing capacity of the soil by inserts of group and reticulated micropiles. In: International Symposium on In-Situ Reinforcement of Soils and Rocks, Paris, 83–89 (1984) Poulos, H.G., Davis, E.H.: Pile Foundation Analysis and Design. Wiley, New York (1980) Singh, S., Heine, E.I.: Underpinning existing footings with micro-piles. In: Proceedings of International Conference on Case Histories in Geotechnical Engineering, St. Louis, vol. 3, pp. 1373–1383 (1984) Stuedlein, A.W., Gibson, M.D., Horvitz, G.E.: Tension and compression micro-pile load tests in gravely sand. In: 16th International Conference on Case Histories in Geotechnial Engineering, Arlington, VA, Paper No 1.12, pp. 1–11 (2008) Wolosick, J.: Macropile load test. In: International Workshop on Micropiles (IWM) Seattle, WA (2003)

Author Index

A Alassal, Mohamed A., 61 B Bathurst, R. J., 73 Belghit, A., 51 Benanane, A., 51 Bouali, Meriem Fakhreddine, 89 Bouassida, Mounir, 89 Bouhamou, N., 51 Bourdim, S. M. A., 51 D Dai, Guo-liang, 18 Damians, I. P., 73 Darwish, Mohamed, 7 E El-Arabi, Dalia, 7 Elmamlouk, Hussein H., 61 El-Menoufy, Ali, 7 F Farid, Ahmed T., 154 Fatahi, Behzad, 40 G Gong, Wei-ming, 18 H Halim, Fady, 7 Hassan, Asmaa M., 61 Henry, Sara, 7

J Josa, A., 73 K Khali, Rana, 7 Khedr, Safwan, 7 L Li, Xiao-juan, 18 Lloret, A., 73 M Mansouri, D. El, 73 Meena, Naveen Kumar, 40 N Nimbalkar, Sanjay, 40 O Oliveira, Ricardo, 106 P Portugal, Alexandre, 106 S Said, Joe, 7 Saleh, Amani, 7 Sêco e Pinto, Pedro S., 106 Soylemez, Safak, 1 Y Yousef, Mostafa A., 154 Z Zhu, Ming-xing, 18

© Springer Nature Switzerland AG 2020 P. Pinto et al. (Eds.): GeoMEast 2019, SUCI, p. 165, 2020. https://doi.org/10.1007/978-3-030-34190-9