Materials Characterisation V: Computational Methods and Experiments (Wit Transactions on Engineering Sciences)

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Materials Characterisation V Computational Methods and Experiments

WIT Press publishes leading books in Science and Technology. Visit our website for new and current list of titles. www.witpress.com

WITeLibrary Home of the Transactions of the Wessex Institute. Papers presented at Materials Characterisation V are archived in the WIT eLibrary in volume 72 of WIT Transactions on Engineering Sciences (ISSN 1743-3533). The WIT eLibrary provides the international scientific community with immediate and permanent access to individual papers presented at WIT conferences. http://library.witpress.com

FIFTH INTERNATIONAL CONFERENCE ON COMPUTATIONAL METHODS AND EXPERIMENTS IN MATERIALS CHARACTERISATION MATERIALS CHARACTERISATION 2011

CONFERENCE CHAIRMEN A.A. Mammoli University of New Mexico, USA C.A. Brebbia Wessex Institute of Technology, UK A. Klemm Glasgow Caledonian University, UK

INTERNATIONAL SCIENTIFIC ADVISORY COMMITTEE G. Badalians Gholikandi A. Bayton A. Galybin H. Huh G. Moriconi P. Prochazka I. Sanchez P. Viot

Organised by Wessex Institute of Technology, UK University of New Mexico, USA Sponsored by WIT Transactions on Engineering Sciences

WIT Transactions Transactions Editor Carlos Brebbia Wessex Institute of Technology Ashurst Lodge, Ashurst Southampton SO40 7AA, UK Email: [email protected]

Editorial Board B Abersek University of Maribor, Slovenia Y N Abousleiman University of Oklahoma,

G Belingardi Politecnico di Torino, Italy R Belmans Katholieke Universiteit Leuven,

P L Aguilar University of Extremadura, Spain K S Al Jabri Sultan Qaboos University, Oman E Alarcon Universidad Politecnica de Madrid,

C D Bertram The University of New South

USA

Spain

A Aldama IMTA, Mexico C Alessandri Universita di Ferrara, Italy D Almorza Gomar University of Cadiz, Spain

B Alzahabi Kettering University, USA J A C Ambrosio IDMEC, Portugal A M Amer Cairo University, Egypt S A Anagnostopoulos University of Patras, Greece

M Andretta Montecatini, Italy E Angelino A.R.P.A. Lombardia, Italy H Antes Technische Universitat Braunschweig, Germany

M A Atherton South Bank University, UK A G Atkins University of Reading, UK D Aubry Ecole Centrale de Paris, France H Azegami Toyohashi University of Technology, Japan

A F M Azevedo University of Porto, Portugal J Baish Bucknell University, USA J M Baldasano Universitat Politecnica de Catalunya, Spain J G Bartzis Institute of Nuclear Technology, Greece A Bejan Duke University, USA M P Bekakos Democritus University of Thrace, Greece

Belgium

Wales, Australia

D E Beskos University of Patras, Greece S K Bhattacharyya Indian Institute of Technology, India

E Blums Latvian Academy of Sciences, Latvia J Boarder Cartref Consulting Systems, UK B Bobee Institut National de la Recherche Scientifique, Canada

H Boileau ESIGEC, France J J Bommer Imperial College London, UK M Bonnet Ecole Polytechnique, France C A Borrego University of Aveiro, Portugal A R Bretones University of Granada, Spain J A Bryant University of Exeter, UK F-G Buchholz Universitat Gesanthochschule Paderborn, Germany

M B Bush The University of Western Australia, Australia

F Butera Politecnico di Milano, Italy J Byrne University of Portsmouth, UK W Cantwell Liverpool University, UK D J Cartwright Bucknell University, USA P G Carydis National Technical University of Athens, Greece

J J Casares Long Universidad de Santiago de Compostela, Spain

M A Celia Princeton University, USA A Chakrabarti Indian Institute of Science, India

A H-D Cheng University of Mississippi, USA

J Chilton University of Lincoln, UK C-L Chiu University of Pittsburgh, USA H Choi Kangnung National University, Korea A Cieslak Technical University of Lodz, Poland

S Clement Transport System Centre, Australia M W Collins Brunel University, UK J J Connor Massachusetts Institute of Technology, USA

M C Constantinou State University of New York at Buffalo, USA

D E Cormack University of Toronto, Canada M Costantino Royal Bank of Scotland, UK D F Cutler Royal Botanic Gardens, UK W Czyczula Krakow University of Technology, Poland

M da Conceicao Cunha University of Coimbra, Portugal

L Dávid Károly Róbert College, Hungary A Davies University of Hertfordshire, UK M Davis Temple University, USA A B de Almeida Instituto Superior Tecnico, Portugal

E R de Arantes e Oliveira Instituto Superior Tecnico, Portugal L De Biase University of Milan, Italy R de Borst Delft University of Technology, Netherlands G De Mey University of Ghent, Belgium A De Montis Universita di Cagliari, Italy A De Naeyer Universiteit Ghent, Belgium W P De Wilde Vrije Universiteit Brussel, Belgium L Debnath University of Texas-Pan American, USA N J Dedios Mimbela Universidad de Cordoba, Spain G Degrande Katholieke Universiteit Leuven, Belgium S del Giudice University of Udine, Italy G Deplano Universita di Cagliari, Italy I Doltsinis University of Stuttgart, Germany M Domaszewski Universite de Technologie de Belfort-Montbeliard, France J Dominguez University of Seville, Spain K Dorow Pacific Northwest National Laboratory, USA W Dover University College London, UK C Dowlen South Bank University, UK

J P du Plessis University of Stellenbosch, South Africa

R Duffell University of Hertfordshire, UK A Ebel University of Cologne, Germany E E Edoutos Democritus University of Thrace, Greece

G K Egan Monash University, Australia K M Elawadly Alexandria University, Egypt K-H Elmer Universitat Hannover, Germany D Elms University of Canterbury, New Zealand M E M El-Sayed Kettering University, USA D M Elsom Oxford Brookes University, UK F Erdogan Lehigh University, USA F P Escrig University of Seville, Spain D J Evans Nottingham Trent University, UK J W Everett Rowan University, USA M Faghri University of Rhode Island, USA R A Falconer Cardiff University, UK M N Fardis University of Patras, Greece P Fedelinski Silesian Technical University, Poland

H J S Fernando Arizona State University, USA

S Finger Carnegie Mellon University, USA J I Frankel University of Tennessee, USA D M Fraser University of Cape Town, South Africa

M J Fritzler University of Calgary, Canada U Gabbert Otto-von-Guericke Universitat Magdeburg, Germany

G Gambolati Universita di Padova, Italy C J Gantes National Technical University of Athens, Greece

L Gaul Universitat Stuttgart, Germany A Genco University of Palermo, Italy N Georgantzis Universitat Jaume I, Spain P Giudici Universita di Pavia, Italy F Gomez Universidad Politecnica de Valencia, Spain

R Gomez Martin University of Granada, Spain

D Goulias University of Maryland, USA K G Goulias Pennsylvania State University, USA

F Grandori Politecnico di Milano, Italy W E Grant Texas A & M University, USA

S Grilli University of Rhode Island, USA

R H J Grimshaw Loughborough University, D Gross Technische Hochschule Darmstadt,

M Karlsson Linkoping University, Sweden T Katayama Doshisha University, Japan K L Katsifarakis Aristotle University of

R Grundmann Technische Universitat

J T Katsikadelis National Technical

A Gualtierotti IDHEAP, Switzerland R C Gupta National University of Singapore,

E Kausel Massachusetts Institute of

UK

Germany

Dresden, Germany

Singapore J M Hale University of Newcastle, UK K Hameyer Katholieke Universiteit Leuven, Belgium C Hanke Danish Technical University, Denmark K Hayami University of Toyko, Japan Y Hayashi Nagoya University, Japan L Haydock Newage International Limited, UK A H Hendrickx Free University of Brussels, Belgium C Herman John Hopkins University, USA S Heslop University of Bristol, UK I Hideaki Nagoya University, Japan D A Hills University of Oxford, UK W F Huebner Southwest Research Institute, USA J A C Humphrey Bucknell University, USA M Y Hussaini Florida State University, USA W Hutchinson Edith Cowan University, Australia T H Hyde University of Nottingham, UK M Iguchi Science University of Tokyo, Japan D B Ingham University of Leeds, UK L Int Panis VITO Expertisecentrum IMS, Belgium N Ishikawa National Defence Academy, Japan J Jaafar UiTm, Malaysia W Jager Technical University of Dresden, Germany Y Jaluria Rutgers University, USA C M Jefferson University of the West of England, UK P R Johnston Griffith University, Australia D R H Jones University of Cambridge, UK N Jones University of Liverpool, UK D Kaliampakos National Technical University of Athens, Greece N Kamiya Nagoya University, Japan D L Karabalis University of Patras, Greece

Thessaloniki, Greece

University of Athens, Greece

Technology, USA

H Kawashima The University of Tokyo, Japan

B A Kazimee Washington State University, USA

S Kim University of Wisconsin-Madison, USA D Kirkland Nicholas Grimshaw & Partners Ltd, UK

E Kita Nagoya University, Japan A S Kobayashi University of Washington, USA

T Kobayashi University of Tokyo, Japan D Koga Saga University, Japan S Kotake University of Tokyo, Japan A N Kounadis National Technical University of Athens, Greece

W B Kratzig Ruhr Universitat Bochum, Germany

T Krauthammer Penn State University, USA C-H Lai University of Greenwich, UK M Langseth Norwegian University of Science and Technology, Norway

B S Larsen Technical University of Denmark, Denmark

F Lattarulo Politecnico di Bari, Italy A Lebedev Moscow State University, Russia L J Leon University of Montreal, Canada D Lewis Mississippi State University, USA S lghobashi University of California Irvine, USA

K-C Lin University of New Brunswick, Canada

A A Liolios Democritus University of Thrace, Greece

S Lomov Katholieke Universiteit Leuven, Belgium

J W S Longhurst University of the West of England, UK

G Loo The University of Auckland, New Zealand

J Lourenco Universidade do Minho, Portugal J E Luco University of California at San Diego, USA

H Lui State Seismological Bureau Harbin, China

C J Lumsden University of Toronto, Canada L Lundqvist Division of Transport and

Location Analysis, Sweden T Lyons Murdoch University, Australia Y-W Mai University of Sydney, Australia M Majowiecki University of Bologna, Italy D Malerba Università degli Studi di Bari, Italy G Manara University of Pisa, Italy B N Mandal Indian Statistical Institute, India Ü Mander University of Tartu, Estonia H A Mang Technische Universitat Wien, Austria G D Manolis Aristotle University of Thessaloniki, Greece W J Mansur COPPE/UFRJ, Brazil N Marchettini University of Siena, Italy J D M Marsh Griffith University, Australia J F Martin-Duque Universidad Complutense, Spain T Matsui Nagoya University, Japan G Mattrisch DaimlerChrysler AG, Germany F M Mazzolani University of Naples “Federico II”, Italy K McManis University of New Orleans, USA A C Mendes Universidade de Beira Interior, Portugal R A Meric Research Institute for Basic Sciences, Turkey J Mikielewicz Polish Academy of Sciences, Poland N Milic-Frayling Microsoft Research Ltd, UK R A W Mines University of Liverpool, UK C A Mitchell University of Sydney, Australia K Miura Kajima Corporation, Japan A Miyamoto Yamaguchi University, Japan T Miyoshi Kobe University, Japan G Molinari University of Genoa, Italy T B Moodie University of Alberta, Canada D B Murray Trinity College Dublin, Ireland G Nakhaeizadeh DaimlerChrysler AG, Germany M B Neace Mercer University, USA D Necsulescu University of Ottawa, Canada F Neumann University of Vienna, Austria S-I Nishida Saga University, Japan

H Nisitani Kyushu Sangyo University, Japan B Notaros University of Massachusetts, USA P O’Donoghue University College Dublin, Ireland

R O O’Neill Oak Ridge National Laboratory, USA

M Ohkusu Kyushu University, Japan G Oliveto Universitá di Catania, Italy R Olsen Camp Dresser & McKee Inc., USA E Oñate Universitat Politecnica de Catalunya, Spain

K Onishi Ibaraki University, Japan P H Oosthuizen Queens University, Canada E L Ortiz Imperial College London, UK E Outa Waseda University, Japan A S Papageorgiou Rensselaer Polytechnic Institute, USA

J Park Seoul National University, Korea G Passerini Universita delle Marche, Italy B C Patten University of Georgia, USA G Pelosi University of Florence, Italy G G Penelis Aristotle University of Thessaloniki, Greece

W Perrie Bedford Institute of Oceanography, Canada

R Pietrabissa Politecnico di Milano, Italy H Pina Instituto Superior Tecnico, Portugal M F Platzer Naval Postgraduate School, USA D Poljak University of Split, Croatia V Popov Wessex Institute of Technology, UK H Power University of Nottingham, UK D Prandle Proudman Oceanographic Laboratory, UK

M Predeleanu University Paris VI, France M R I Purvis University of Portsmouth, UK I S Putra Institute of Technology Bandung, Indonesia

Y A Pykh Russian Academy of Sciences, Russia

F Rachidi EMC Group, Switzerland M Rahman Dalhousie University, Canada K R Rajagopal Texas A & M University, USA T Rang Tallinn Technical University, Estonia J Rao Case Western Reserve University, USA A M Reinhorn State University of New York at Buffalo, USA

A D Rey McGill University, Canada

D N Riahi University of Illinois at Urbana-

Champaign, USA B Ribas Spanish National Centre for Environmental Health, Spain K Richter Graz University of Technology, Austria S Rinaldi Politecnico di Milano, Italy F Robuste Universitat Politecnica de Catalunya, Spain J Roddick Flinders University, Australia A C Rodrigues Universidade Nova de Lisboa, Portugal F Rodrigues Poly Institute of Porto, Portugal C W Roeder University of Washington, USA J M Roesset Texas A & M University, USA W Roetzel Universitaet der Bundeswehr Hamburg, Germany V Roje University of Split, Croatia R Rosset Laboratoire d’Aerologie, France J L Rubio Centro de Investigaciones sobre Desertificacion, Spain T J Rudolphi Iowa State University, USA S Russenchuck Magnet Group, Switzerland H Ryssel Fraunhofer Institut Integrierte Schaltungen, Germany S G Saad American University in Cairo, Egypt M Saiidi University of Nevada-Reno, USA R San Jose Technical University of Madrid, Spain F J Sanchez-Sesma Instituto Mexicano del Petroleo, Mexico B Sarler Nova Gorica Polytechnic, Slovenia S A Savidis Technische Universitat Berlin, Germany A Savini Universita de Pavia, Italy G Schmid Ruhr-Universitat Bochum, Germany R Schmidt RWTH Aachen, Germany B Scholtes Universitaet of Kassel, Germany W Schreiber University of Alabama, USA A P S Selvadurai McGill University, Canada J J Sendra University of Seville, Spain J J Sharp Memorial University of Newfoundland, Canada Q Shen Massachusetts Institute of Technology, USA X Shixiong Fudan University, China G C Sih Lehigh University, USA L C Simoes University of Coimbra, Portugal

A C Singhal Arizona State University, USA P Skerget University of Maribor, Slovenia J Sladek Slovak Academy of Sciences, Slovakia

V Sladek Slovak Academy of Sciences, Slovakia

A C M Sousa University of New Brunswick, Canada

H Sozer Illinois Institute of Technology, USA D B Spalding CHAM, UK P D Spanos Rice University, USA T Speck Albert-Ludwigs-Universitaet Freiburg, Germany

C C Spyrakos National Technical University of Athens, Greece

I V Stangeeva St Petersburg University, Russia

J Stasiek Technical University of Gdansk, Poland

G E Swaters University of Alberta, Canada S Syngellakis University of Southampton, UK J Szmyd University of Mining and Metallurgy, Poland

S T Tadano Hokkaido University, Japan H Takemiya Okayama University, Japan I Takewaki Kyoto University, Japan C-L Tan Carleton University, Canada E Taniguchi Kyoto University, Japan S Tanimura Aichi University of Technology, Japan

J L Tassoulas University of Texas at Austin, USA

M A P Taylor University of South Australia, Australia

A Terranova Politecnico di Milano, Italy A G Tijhuis Technische Universiteit Eindhoven, Netherlands

T Tirabassi Institute FISBAT-CNR, Italy S Tkachenko Otto-von-Guericke-University, Germany

N Tosaka Nihon University, Japan T Tran-Cong University of Southern Queensland, Australia

R Tremblay Ecole Polytechnique, Canada I Tsukrov University of New Hampshire, USA R Turra CINECA Interuniversity Computing Centre, Italy

S G Tushinski Moscow State University, Russia

J-L Uso Universitat Jaume I, Spain E Van den Bulck Katholieke Universiteit

Z-Y Yan Peking University, China S Yanniotis Agricultural University of Athens,

D Van den Poel Ghent University, Belgium R van der Heijden Radboud University,

A Yeh University of Hong Kong, China J Yoon Old Dominion University, USA K Yoshizato Hiroshima University, Japan T X Yu Hong Kong University of Science &

Leuven, Belgium

Netherlands

R van Duin Delft University of Technology, Netherlands

Greece

Technology, Hong Kong

P Vas University of Aberdeen, UK R Verhoeven Ghent University, Belgium A Viguri Universitat Jaume I, Spain Y Villacampa Esteve Universidad de

M Zador Technical University of Budapest,

F F V Vincent University of Bath, UK S Walker Imperial College, UK G Walters University of Exeter, UK B Weiss University of Vienna, Austria H Westphal University of Magdeburg,

R Zarnic University of Ljubljana, Slovenia G Zharkova Institute of Theoretical and

Alicante, Spain

Germany

J R Whiteman Brunel University, UK

Hungary

K Zakrzewski Politechnika Lodzka, Poland M Zamir University of Western Ontario, Canada

Applied Mechanics, Russia

N Zhong Maebashi Institute of Technology, Japan

H G Zimmermann Siemens AG, Germany

Materials Characterisation V Computational Methods and Experiments

Editors A.A. Mammoli University of New Mexico, USA C.A. Brebbia Wessex Institute of Technology, UK A. Klemm Glasgow Caledonian University, UK

Editors: A.A. Mammoli University of New Mexico, USA C.A. Brebbia Wessex Institute of Technology, UK A. Klemm Glasgow Caledonian University, UK Published by WIT Press Ashurst Lodge, Ashurst, Southampton, SO40 7AA, UK Tel: 44 (0) 238 029 3223; Fax: 44 (0) 238 029 2853 E-Mail: [email protected] http://www.witpress.com For USA, Canada and Mexico Computational Mechanics Inc 25 Bridge Street, Billerica, MA 01821, USA Tel: 978 667 5841; Fax: 978 667 7582 E-Mail: [email protected] http://www.witpress.com British Library Cataloguing-in-Publication Data A Catalogue record for this book is available from the British Library

ISBN: 978-1-84564-538-0 ISSN: 1746-4471 (print) ISSN: 1743-3533 (on-line) The texts of the papers in this volume were set individually by the authors or under their supervision. Only minor corrections to the text may have been carried out by the publisher. No responsibility is assumed by the Publisher, the Editors and Authors for any injury and/ or damage to persons or property as a matter of products liability, negligence or otherwise, or from any use or operation of any methods, products, instructions or ideas contained in the material herein. © WIT Press 2011 Printed in Great Britain by Martins the Printers. All rights reserved. No part of this publication may be reproduced, stored in a retrieval system, or transmitted in any form or by any means, electronic, mechanical, photocopying, recording, or otherwise, without the prior written permission of the Publisher.

Preface

The increasing demands for high-quality products from both industry and consumers are the driving force for the rapid developments in materials science and engineering. In principle materials science involves relating the desired properties and relative performance of a material to its microstructural features through characterization. The major determinants of the structure of a material and hence its properties are its constituent chemical elements and the way a material has been processed into its final form. Over the years, a variety of experimental techniques have been developed for characterizing the physical and chemical properties of materials. Unfortunately due to a number of simplifying assumptions and limitations on the use of individual methods, it is not often possible to describe in a qualitative, reliable way the microstructural features of many materials. Triangulation of different experimental methods as well as computer simulations may become essential to achieve a thorough, comprehensive analysis. Simulations can contribute to the understanding of the phenomena and to provide a good basis for the development of durable materials and components which can withstand ambient and extreme environmental conditions. The way forward in material characterisation is to develop new experimental techniques or apply existing methodologies adopted from other related disciplines. A very wide range of materials, starting with metals through polymers, semiconductors to composites, necessitates a whole spectrum of experimental techniques and numerical models, which are specific for material types. Some of these well established methodologies could potentially find applications in new fields. In this context a multidisciplinary approach in material characterisation and the exchange of original ideas is indispensible. The aim of the International Conference on Computational Methods and Experiments in Materials Characterisation held in Kos, Greece, was therefore to facilitate such interdisciplinary interactions within the research community. The

resulting conference book has been arranged in several chapters addressing various experimental and numerical methods. The wide range of topics covers mechanical characterisation and testing, corrosion problems and thermal analysis as well as recycled materials, nano-composites and energy materials. The editors would like to express their gratitude to all authors without whose involvement this book could not have been produced. We wish to aknowledge the valuable input of the members of the Scientific Advisory Committee in attracting and selecting many high quality contributions. We trust that this book will present some innovative ideas and will facilitate further developments in materials science. The Editors, Kos, Greece 2011

Contents Section 1: Micro and nano characterisation of cementitious materials (Special session organised by A. J. Klemm) Application of positron annihilation lifetime spectroscopy to nano-characterisation of polymer-modified mortars P. Guagliardo, A. J. Klemm, S. N. Samarin & J. F. Williams ............................. 3 Multi-technique investigation of calcium hydroxide crystals at the concrete surface E. Gueit, E. Darque-Ceretti, P. Tintillier & M. Horgnies ................................. 15 Characterization of the influence of the casting mould on the surface properties of concrete and on the adhesion of a protective coating M. Horgnies, P. Willieme, O. Gabet, S. Lombard & M. Dykman...................... 27 Section 2: Nano-materials HRTEM techniques applied to nanocrystal modeling: towards an “atom-by-atom” description D. G. Stroppa, L. A. Montoro, E. R. Leite & A. J. Ramirez ............................... 41 Ca(OH)2 nanoparticle characterization: microscopic investigation of their application on natural stones V. Daniele & G. Taglieri ................................................................................... 55 Nanocarbon composite materials with optical response on radioactive waste M. Vantsyan, G. Popova, E. Karpuzova, M. Bobrov, O. Plaksin & E. Dabek ........................................................................................................ 67

Section 3: Corrosion problems Evaluation of the fretting corrosion mechanisms on the head-cone interface of hip prostheses I. Caminha, C. R. M. Roesler, H. Keide, C. Barbosa, I. Abud & J. L. Nascimento ............................................................................................ 77 Improving corrosion performance by surface patterning M. Bigdeli Karimi, V. Stoilov & D. O. Northwood ............................................ 85 Material characterisation to understand various modes of corrosion failures in large military vehicles of historical importance A. Saeed, Z. Khan, N. Garland & R. Smith........................................................ 95 Section 4: Computational models and experiments A multi-factor interaction model (MFIM) for damage initiation and progression C. C. Chamis.................................................................................................... 109 Analytical solution of a two-dimensional elastostatic problem of functionally graded materials via the Airy stress function H. Sakurai........................................................................................................ 119 Moment curvature analysis of concrete flexural members confined with CFRP grids A. Michael & P. Christou ................................................................................ 131 Application of effective media theory in the characterization of the hygrothermal performance of masonry Z. Pavlík, E. Vejmelková, L. Fiala, M. Pavlíková & R. Černý......................... 143 3D FIB reconstruction and characterisation of a SOFC electrode S. Chupin, N. Vivet, D. Rochais & E. Bruneton............................................... 155 Modelling of load transfer between porous matrix and short fibres in ceramic matrix composites J. G. P. Silva, D. Hotza, R. Janssen & H. A. Al-Qureshi................................. 165 Modeling aspects concerning the axial behavior of RC columns H. O. Koksal, T. Turgay, C. Karakoç & S. Ayçenk.......................................... 175

Section 5: Innovative experiments Surface characterization of eucalyptus and ash wood veneers by XPS, TOF-SIMS, optic profilometry and contact angle measurements G. Vázquez, R. Ríos, M. S. Freire, G. Antorrena & J. González-Álvarez........ 187 Interface resistances in heat and moisture transport: semi-scale experimental analysis Z. Pavlík, J. Mihulka, J. Žumár, M. Pavlíková & R. Černý ............................. 199 Section 6: Mechanical characterisation and testing Tension/compression test of auto-body steel sheets with the variation of the pre-strain and the strain rate G. H. Bae & H. Huh ........................................................................................ 213 Definition of averaged elastic-plastic characteristics of sandwich panel structures I. I. Zakirov, V. N. Paimushin & I. M. Zakirov ................................................ 227 Hot deformation and mechanical properties of P/M Al special M. Tercelj, P. Cvahte, I. Perus & G. Kugler ................................................... 239 Coarsening kinetics of the bimodal γ′ distribution in DS GTD111TM superalloy V. S. K. G. Kelekanjeri, S. K. Sondhi, T. Vishwanath, F. Mastromatteo & B. Dasan ...................................................................................................... 251 Effect of the elastomer stiffness and coupling agents on rheological properties of magnetorheological elastomers A. Boczkowska & S. F. Awietjan...................................................................... 263 Optimization of magnetoelastic properties of pure nickel by means of heat treatments A. L. Morales, A. J. Nieto, J. M. Chicharro, P. Pintado, G. P. Rodríguez & G. Herranz................................................................................................... 275 Nanomechanical structure-property relations of dynamically loaded reactive powder concrete P. G. Allison, R. D. Moser, M. Q. Chandler, T. S. Rushing, B. A. Williams & T. K. Cummins ............................................................................................. 287 Dynamic strength of concrete under multiaxial compressive loading Y. P. Song & H. L. Wang ................................................................................. 299

Modelling and simulation of the rutting resistance of bituminous mixes: experimental and stochastic approaches A. E. Ouni, A. Dony & J. Colin........................................................................ 307 Laboratory tests on the cleanliness of soil materials used as subgrades in pavement structures A. Athanasopoulou & G. Kollaros................................................................... 315 Use of additives to improve the engineering properties of swelling soils in Thrace, Northern Greece A. Athanasopoulou & G. Kollaros................................................................... 327 Characteristics of a bolted joint with a shape memory alloy stud N. Ould-Brahim, A.-H. Bouzid & V. Brailovski............................................... 339 Section 7: Thermal analysis Experimental validation of a thermal model of adhesively bonded scarf repairs for CFRP composite materials incorporating cure kinetics C. C. N. Bestley, S. G. R. Brown & S. M. Alston ............................................. 351 Computational and experimental characterization of building envelopes based on autoclaved aerated concrete V. Kočí, J. Výborný & R. Černý....................................................................... 363 Section 8: Recycled materials Quantitative description of the morphology of polyurethane nanocomposites for medical applications J. Ryszkowska & B. Waśniewski ...................................................................... 377 Description methods of the properties of composites from oxybiodegradable foil waste and wood J. Ryszkowska & K. Sałasińska........................................................................ 387 The effect of slag composition on recycling of “OFHC” through the “ESCM” process S. Ketabchi, F. K. Ahadi, K. Hanaee & S. H. Alhoseini .................................. 397 Author Index .................................................................................................. 407

Section 1 Micro and nano characterisation of cementitious materials (Special session organised by A. J. Klemm)

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Materials Characterisation V

3

Application of positron annihilation lifetime spectroscopy to nano-characterisation of polymer-modified mortars P. Guagliardo1, A. J. Klemm2, S. N. Samarin1 & J. F. Williams1 1

ARC Centre of Excellence for Antimatter-Matter Studies, School of Physics, University of Western Australia, Australia 2 School of Built and Natural Environment, Glasgow Caledonian University, UK

Abstract Positron annihilation lifetime spectroscopy (PALS) has been applied to study the microstructural features of immature cement mortars. Two types of cement mortars containing superabsorbent polymers (SAPs) were studied, in addition to Ordinary Portland Cement (OPC). The ortho-positronium lifetimes for all samples were in the range of 1.70-1.73 ns, values that are close to that of free water (1.7 ns) and hence suggest the presence of water-filled pores. Periodic lifetime measurements showed that the intensity of this component decreased slightly over a period of four weeks, indicating water loss associated with the curing process, evaporation or a combination of the two. Keywords: cement mortar, superabsorbent polymer, positron annihilation lifetime spectroscopy, positronium, porosity, hydration, curing.

1 Introduction Cement is a material of immense practical importance, but in spite of its almost ubiquitous use, its microstructural characterisation still proves to be problematic. The complexity of the problem is enhanced when the composition of cement is modified through the use of auxiliary agents. Ordinary Portland cement is a combination of gypsum (CS*H2) and clinker. Gypsum acts to prevent rapid setting of the cement paste while clinker is the hydraulic binder. The major constituents of clinker are tricalcium silicate (C3S), dicalcium silicate (C2S), WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line) doi:10.2495/MC110011

4 Materials Characterisation V tricalcium aluminate and tetracalcium alumina-ferrite (C4AF); C = CaO, S = SiO2, A = Al2O3, F = Fe2O3, S* = SO3 and H = H2O (abbreviations commonly used in industrial nomenclature) [1, 2]. The presence of water among the active compounds of cement leads to a rearrangement of their structure and the initiation of the hardening process. The initial hardening is caused by the hydration of C3S, which forms a gel-like silicate and calcium hydrate phase referred to as C-S-H gel. These particles ultimately crystallise and bind together the particles of sand or stone into a hard mass. Other hydrates of the complexes described above are also formed, although the reaction rates may differ considerably. The final product is a hardened cement paste which is a mixture of unreacted cement particles, hydration products and pores. Two types of pores are usually distinguished capillary pores, which comprise the water- or air-filled spaces between the hydrates, with sizes in the range of 10-1000 nm, and smaller pores referred to as gel pores, which are contained in the amorphous hydrate phase. Gel and capillary pores form a continuous network of pores throughout the material [3, 4]. The mechanical properties of cementitious materials are heavily influenced by porosity, with the volume and size distribution of pores controlling both the strength and durability of the material [5]. A detailed characterisation of the pore structure and the factors that affect it are thus crucial to advancing the design of these materials; however, classical porosimetry methods such as mercury intrusion porosimetry (MIP) and gas adsorption cannot always be relied upon to extract information on pores smaller than a few nanometres [6]. In this regard, positron annihilation lifetime spectroscopy (PALS) has considerable advantages over these classical methods. PALS is sensitive to both open and closed porosity, for example, whereas isolated pores are invisible to the aforementioned techniques. In addition, PALS is sensitive to pores in the size range of 0.3-30 nm (that is, in the size range of gel pores), and both pore sizes and relative concentrations can be measured (refer to the following section for a detailed explanation) [7]. In this work, PALS was used to study the features of the porosity of three cement mortars. Two types of cement mortars containing superabsorbent polymers (SAPs) were studied, in addition to a sample composed only of OPC. SAPs are cross-linked networks of hydrophilic polymers with a high capacity for water uptake - they can absorb and retain up to 500 times their weight in water. This makes them ideal for use in water-absorbing applications such as absorbent medical dressings and controlled release media [8]. During the mixing process the polymers absorb the pore solution immediately after their addition to the mortar, reaching saturation within minutes [9, 10]. They then swell to form spherical cavities filled with water. At later stages of the hydration process the water is released to the concrete matrix and the cavities remain as empty pores. For mortars with a low water-to-cement ratio, it is possible to replace part of the irregular capillary pores with larger spherical pores formed by saturated polymers. The dispersion and size of these pores can be estimated by the material attributes of polymers.

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5

The samples were studied with PALS over a period of 4 weeks with measurements starting when the samples were aged 21 days, in order to elucidate the changes in nano-porosity caused by the addition of SAPs as well as the changes arising from the curing/hydration process. In addition, samples aged at 3 and 6 weeks were studied with MIP to determine changes in their pore distributions and bulk densities. MIP can provide information on larger pores (from 100 nm to 100 µm) to which PALS is not sensitive.

2 Positron annihilation background Positron annihilation techniques have been applied to study nano-porosity in a wide range of materials, such as zeolite, silicates and polymers [7]. In positron annihilation lifetime spectroscopy (PALS), positrons from a 22Na radioisotope are injected into the sample under study and a positron lifetime spectrum is recorded [11]. As a 1.274 MeV gamma-quantum is emitted during the decay of 22 Na almost simultaneously with a positron, the lifetime is defined as the time between detection of the 1.274 MeV photon and the subsequent 511 keV photons created from the positron’s annihilation with an electron. The measured spectrum is a histogram of the time periods between these two events. Ideally, it is a sum of decaying exponentials of the form

∑

/

(1)

positron state with lifetime . An where is the intensity or weighting of the experimentally obtained spectrum differs from this form however, in that it is convolved with the time response function of the apparatus, usually approximated by a sum of Gaussians. Porosimetry is made possible by exploiting the phenomenon of positronium (Ps) formation, Ps being the hydrogen-like bound state of a positron and an electron [12]. In brief, positrons penetrate the material under investigation and rapidly lose energy, predominately through ionising collisions. Ps formation can then take place as a reaction between the positron and a secondary electron produced during the positrons’ thermalisation (these secondary electrons are often termed spur electrons [13]). Positronium is formed in two spin states with dramatically different annihilation characteristics. The singlet state (with antiparallel orientation of spins) is termed para-positronium (p-Ps); it has a vacuum lifetime of 125 ps and decays via the emission of two gamma quanta. The triplet state (with parallel spins of electron and positron) is referred to as orthopositronium (o-Ps) and has a much longer vacuum lifetime of 142 ns, decaying via three gamma emission. Due to its short lifetime, p-Ps is not significantly perturbed by the material; however, o-Ps, because of its intrinsically longer lifetime, interacts strongly with the material and its pore structure. In the presence of matter, o-Ps can decay into two gammas via a process known as pick-off annihilation - the positron in o-Ps annihilates with an electron of opposite spin in the material via a two-gamma process. As a result of pick-off WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)

6 Materials Characterisation V annihilation, the o-Ps lifetime in matter is significantly shorter than its vacuum value (and can be as low as ~ 1 ns) [7]. The basis for studying porosity is that o-Ps tends to localise in regions of low density, such as pores. This is because in the bulk of the material the positron and the electron in o-Ps experience repulsive Coulomb and exchange forces respectively. As a result, most of these materials (porous insulators) have a negative work function for positronium [7]. Porosimetry is possible because the annihilation rate (the inverse of the lifetime) in the pore is a function of the pore size – smaller pores have a shorter lifetime compared to large pores, where the overlap of the Ps wave function with the pore walls is reduced. A semi-empirical quantum mechanical relation has been developed by Tao [14] and Eldrup et al. [15] to relate the o-Ps lifetimes to pore radii: 1 τ

= 2 1-

R R+ΔR

+

1 2π

sin

2πR

(2)

R+ΔR

where  is the lifetime, R+R is the pore radius and R is empirically determined to be 0.16-0.17 nm. This model is applicable to sub-nanometre pores of a spherical geometry. Gidley et al. [16, 17] extended this model to include larger pore sizes and varying pore geometries (cylinders, cubes, channels and sheets).

3 Experimental 3.1 Materials and mixes For the purpose of this research Portland cement (BS EN 197-1 CEM II/B-V 32,5) was mixed at 1:1 ratio with fine sand (the vast majority of particles were distributed below 0.6 mm). Throughout the investigation the total water-tocement ratio of 0.45 was maintained. The mix compositions are presented in Table 1. The pastes were shaped into cylinders with diameters of about 25 mm and thicknesses in the range of 5-10 mm. Table 1:

Composition of cement mortars.

Sample designation

OPC

SAP-A

SAP-B

Mix code

R

A

B

(Water/Cement)total

0.45

0.45

0.45

(Water/Cement)effective

0.45

0.425

0.438

Sand/Cement

1

1

1

SAP content [%] (by cement weight)

0

0.25

0.25



7

The SAPs used in this study were cross-linked polymers provided by BASF. SAP-A is a copolymer of acrylamide and acrylic acid and SAP-B is a polymer based on acrylic acid. The products had absorption capacities of 200-250 ml/g in demineralised water, though the absorption in mortar depended on the product and was approximately 10 mL/g for SAP-A and 5 mL/g for SAP-B. Both materials were prepared by grinding and screening to sizes of 63-125 µm, but there was also a minor (less than 10%) content of finer particles. The SAP particle shapes were irregular and their sizes after initial absorption were in the range of 135-270 µm for SAP-A and 105-210 µm for SAP-B. 3.2 Positron lifetime and mercury intrusion measurements For positron lifetime measurements, approximately 30 μCi of 22NaCl was deposited on 7μm Kapton foil and covered with an identical foil. The edges of this foil-sandwich were then sealed, and this source foil was placed between two identical pieces of sample. The positron lifetimes were measured with a fast-fast coincidence system. The gamma-ray detectors consist of a truncated cone (31.8 mm diameter tapering to 12.7 mm with a height of 12.7 mm) BC418 scintillator coupled to a Burle 8850 photomultiplier tube. The time resolution of the system is approximately 220 ps, as determined from analysis of a spectrum of highpurity annealed nickel. The spectra comprise of at least 2 million counts and have been analysed using PAScual version 1.3.0 [18]. Mercury intrusion porosimetry (MIP) has been carried out with the use of a Porosimeter Autopore IV 9500 by Micromeritics, with a pressure range up to 60000 psi for all samples at ages 3 and 6 weeks.

4 Results and discussion In Table 2 the fitted lifetimes and intensities for all samples are given at four different ages. In each case, the lifetime spectra could be fitted with three discrete components. In porous media a number of annihilation processes are possible because multiple positron and positronium states exist. It is likely that the two shorter lifetime components (1 and 2) in Table 2 are the averages of several different annihilation modes. These could include positron annihilation (of both free and trapped positrons), p-Ps self-annihilation and o-Ps pick-off annihilation in the bulk of the material, and annihilations in the source foil. Nanosecond lifetime components (3) on the other hand are typically associated with o-Ps pick-off annihilation in nano-pores. For all of the cement samples the average 3 over the course of the measurements was in the range of 1.70-1.73 ns. The presence of a nanosecond component is often an indicator of porosity or free volume, and these values would correspond to pores having diameters in the range of 0.51-0.52 nm (using the TE model, equation 2). However, these lifetimes are very close to that observed in pure water, approximately 1.7 ns [19]. Given that cement pastes are known to contain a significant fraction of water, and that the samples were tested at a relatively young age (testing began at 21 days after mixing), it seems likely that these WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)

8 Materials Characterisation V nanosecond components are associated with annihilation in water contained in capillary pores of the cement and hence cannot be directly related to a pore size. Consolati and Quasso [20, 21], Consolati et al. [22] and Salgueiro et al. [23] observed similar lifetimes in Portland cement pastes and attributed this component to water-filled pores.

OPC

SAP-B

SAP-A

Table 2: Age (weeks) 3 4 5 6 3 4 5 6 3 4 5 6

Lifetimes and intensities; standard deviations are given in brackets. 1 (ps)

2 (ps)

218 (9) 227 (9) 242 (7) 233 (8) 219 (5) 229 (5) 221 (6) 237 (6) 224 (6) 239 (5) 238 (5) 212 (9)

410 (10) 410 (10) 430 (10) 420 (10) 409 (7) 421 (9) 409 (8) 420 (10) 402 (8) 430 (10 420 (10) 390 (10)

3 (ns)

I1 (%)

I2 (%)

I3 (%)

2

1.67 (0.02) 1.67 (0.03) 1.73 (0.03) 1.72 (0.03) 1.73 (0.02) 1.76 (0.03) 1.71 (0.03) 1.63 (0.03) 1.71 (0.02) 1.67 (0.02) 1.67 (0.03) 1.73 (0.03)

43 (4) 46 (5) 55 (4) 50 (4) 45 (2) 49 (3) 46 (3) 53 (4) 47 (3) 57 (3) 53 (3) 43 (3)

50 (3) 48 (5) 40 (4) 44 (4) 49 (2) 46 (3) 50 (3) 43 (4) 48 (3) 39 (2) 42 (3) 53 (4)

5.9 (0.1) 5.7 (0.2) 5.2 (0.2) 5.2 (0.2) 5.0 (0.1) 4.0 (0.1) 4.3 (0.1) 4.2 (0.2) 4.6 (0.1) 4.3 (0.1) 4.2 (0.1) 4.3 (0.1)

0.93 1.04 0.99 0.99 1.15 1.08 1.04 1.08 1.01 1.01 1.09 1.07

In Figure 1, o-Ps lifetimes (3) and intensities (I3) are plotted as a function of sample age. Lifetime spectra were recorded every 7 days for a 4 week period. Over the course of the measurements, I3 for the cement sample containing SAPA is somewhat higher than for the cement containing SAP-B and the pure cement sample. If these lifetimes are indicative of water-filled pores then this suggests a higher concentration of pores in the sample containing SAP-A. As the results for SAP-B are closer to those of pure cement, this suggests that the presence of the polymer does not significantly alter the pore concentration. For the SAP-A sample, a gradual decrease in I3 is observed over weeks 3-5, whereas for the SAP-B and pure cement samples there is a decrease after the initial measurement (weeks 3-4), and then I3 is relatively constant. While these changes are small, they may be indicative of various stages of the hydration process, as the decrease in I3 is consistent with the reduction in total porosity that is known to occur as hydration progresses [3]. The gradual decline seen in the SAP-A sample suggests that this polymer releases its water more slowly compared to SAP-B. At the onset of the measurement (age 21 days) the hydration process is likely to be still underway, with both di- and tri-calcium silicates reacting with water to produce a network of calcium silicate hydrates. Water could then be consumed in the hydration process for at least the next two weeks. However, during this time it is difficult to say if the loss of water seen by the decrease in I3 is also associated with evaporation. As the samples lose water one might expect the lifetimes to increase as there would be a transition from filled to empty pores.



9

The fact that this does not occur suggests that pores collapse with water loss (the volume of capillary pores is known to decrease with hydration [3]). These results are consistent with the work of Consolati and Quasso [20, 21] who observed a systematic decrease in I3 with sample age, but they are in contrast to the work of Myllylä and Karras [24] where the opposite trend was seen. It should be noted however that a direct comparison with these studies is not possible due to differing experimental conditions (namely the composition and age of samples).

o-Ps intensity, I3 (%)

6.5 6.0 5.5 5.0 4.5 4.0 3.5 3

4

5

6

Sample age (weeks)

o-Ps lifetime, 3 (ns)

2.2 2.0 1.8 1.6 1.4 1.2 3

4

5

6

Sample age (weeks) Figure 1:

o-Ps intensities (top) and lifetimes (bottom) vs. sample age for SAP-A (circles), SAP-B (triangles) and Ordinary Portland Cement (squares).

In addition to PALS, MIP measurements were carried out on samples aged 3 and 6 weeks; the resulting pore distributions are shown in Figure 2. It is apparent that the samples contain a wide distribution of pores, with the majority of pores WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)

10 Materials Characterisation V

0.25 0.2 0.15 0.1 0.05 0 1000000 100000

10000

1000

100

10

Log differential intrusion (mL/g)

in the size range of 10-100 nm. The MIP results are in qualitative agreement with the PALS data in the sense that the sample containing SAP-B displays similar behaviour to pure cement - at 6 weeks their pore distributions almost coincide and the total porosity for both samples has decreased with aging. This decrease is also consistent with the reduction in I3 seen in the PALS data.

1

0.25 0.2 0.15 0.1 0.05 0 1000000 100000

10000

1000

100

10

1

Log differential intrusion (mL/g)

Pore size diameter (nm)

Pore size diameter (nm) Figure 2:

Pore size distributions determined by MIP at age 3 weeks (top) and 6 weeks (bottom); OPC – diamonds, SAP-A – squares, SAP-B – triangles.

Conversely, the SAP-A sample shows a significant increase in the pore concentration within the 10-100 nm range at 6 weeks. This is in contrast to the decrease in I3 seen in PALS, although PALS may not be sensitive to the porosity in this size range. In spite of the differing trend, both PALS and MIP show that WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


11

the presence of SAP-A significantly alters the hydration process of the cement mortar. The bulk densities of the samples were also determined by MIP measurements (Table 3). It was found that the density of all samples increased with aging time. Again, the changes in density for the pure cement and SAP-B samples are similar. The change in density for SAP-A is much smaller, which is consistent with the slower rate of change of I3. Table 3: Sample OPC SAP-A SAP-B

Bulk densities determined by MIP at 3 and 6 weeks. Bulk density at 3 weeks (g/mL) 1.83 1.82 1.84

Bulk density at 6 weeks (g/mL) 1.89 1.86 1.91

In order to gain further insight into the nature of the porosity, the samples were heated at 116°C for 4 hours under vacuum conditions (10-5 Torr) and lifetime measurements were repeated. Lifetime spectra were recorded directly after heating and the samples were sealed in a plastic bag containing a desiccant (to prevent water absorption) during the measurement. The resulting lifetimes and intensities are given in Table 4 and the corresponding pore sizes are given in Table 5. At least three types of water can be distinguished in a hydrating cement paste. Chemically-bound water is directly incorporated into the structure of the hydration products, physically-bound water is absorbed on the surfaces of cement particles and reaction products, and there is also “free” water contained in the capillary and gel pores [25]. The heat treatment will have removed a significant fraction of the latter two sources of water; however, much higher temperatures (above 950°C) are required to remove chemically-bound water. Table 4:

Lifetimes and intensities obtained after heat treatment.

1 (ps) 2 (ps) 3 (ns) 4 (ns) I1 (%) I2 (%) I3 (%) I4 (%) 2

OPC 263 (2) 536 (7) 4.3 (0.1) 75 (2) 24 (1) 1.3 (0.1) 1.13

Sample SAP-A 214 (9) 380 (9) 1.32 (0.07) 12.6 (0.7) 36 (8) 60 (7) 3.3 (0.4) 1.0 (0.1) 1.05

SAP-B 219 (4) 413 (5) 1.62 (0.05) 14.9 (0.6) 44 (2) 52 (2) 2.5 (0.1) 0.70 (0.01) 1.01


12 Materials Characterisation V After heat treatment, 3 for OPC increases to 4.3 ns. This corresponds to a gel pore diameter of 0.86 nm using the TE model (equation 2). However, the intensity of this component decreases by a factor of about 3.3; this could suggest that a significant fraction of the pores collapse upon water removal. There is also a significant redistribution of intensity from I2 to I1 after heat treatment. These are the shorter components associated with multiple positron states in the material. The redistribution of intensity from the longer component (2) to the shorter one (1) indicates that the heat treatment has reduced the number of positron traps in the bulk material. This could be associated with an increase in the bulk density after heating. Table 5:

Pore diameters, d3 and d4, calculated from o-Ps lifetimes, and total o-Ps intensities (Io-Ps).

d3 (nm) d4 (nm) Io-Ps before (%) Io-Ps after (%)

OPC 0.86 4.3 1.3

Sample SAP-A 0.42 1.45 5.2 4.3

SAP-B 0.50 1.56 4.2 3.2

For the cement samples containing SAPs an additional low intensity, long lifetime component (4) appears in the spectra after heat treatment (and due to the presence of this long component it was necessary to fix the background to the left of the main coincidence peak to the average value in the analysis). 4 may be associated with pores formed by the polymers which have replaced some of the ordinary pore structure, and its emergence shows that the presence of SAP significantly alters the structure of the cement paste. For the sample containing SAP-A, 3 = 1.32 ns and 4 = 12.6 ns, with respective intensities of 3.3% and 1.0%. These lifetimes would correspond to pores having diameters of 0.42 and 1.45 nm respectively. There is also a redistribution of the weighting of the shorter components (1 and 2); however this is not as pronounced as for ordinary Portland cement. For the sample containing SAP-B, 3 = 1.62 ns and 4 = 14.9 ns, with intensities of 2.5% and 0.7% respectively. These lifetimes correspond to pores having diameters of 0.42 and 1.45 nm respectively. The value of 3 is comparable to that obtained before heat treatment (which was similar to the o-Ps lifetime in water). This could suggest that the heat treatment did not result in the complete removal of free water. For both SAP-A and SAP-B the total o-Ps intensity (I3 + I4) decreases after heat treatment, signalling a reduction in porosity; however, the reduction is larger for the ordinary Portland cement sample.



13

5 Conclusion This study has shown that PALS can be used to monitor the hydration process in cement samples by using the o-Ps intensity as an indicator of the concentration of water-filled pores. In addition, PALS has shown that the addition of a SAP to the cement mixture alters the structure and hydration process of cement. The addition of SAP-A caused the hydration process to progress more gradually compared to OPC; however, the addition of SAP-B gave comparable results to that of OPC. Both PALS and MIP showed that the presence of SAP-A has a significant effect on the hydration process. In addition, MIP showed that the concentration of pores in the 10-100 nm range increases with aging for the SAP-A sample. For the samples with SAP added, heat treatment under evacuation (resulting in the removal of free water) gave rise to an additional long lifetime component and a higher overall o-Ps intensity compared to OPC. This shows that the SAP had modified the pore distribution, replacing some of the regular pore distribution with larger pores formed by saturated polymers.

References [1] [2] [3] [4] [5] [6] [7] [8] [9]

[10]

[11] [12] [13] [14]

Bogue, R. H. The Chemistry of Portland Cement. (Reinhold, 1955). Lea, F. M. The Chemistry of Cement and Concrete. (Arnold, 1970). Neville, A. M. Properties of Concrete. 4th edn, (Longman, 1995). Taylor, H. F. W. Cement Chemistry. (Academic, 1990). Ghosh, S. N. Advances in cement technology (Pergamon, Oxford, 1983). Gregg, S. J. & Sing, K. S. W. Adsorption, Surface Area and Porosity. (Academic, 1982). Jean, Y. C., Mallon, P. E. & Schrader, D. M. Principles and applications of positron and positronium chemistry (World Scientific, 2003). Buchholz, F. L. & Graham, A. T. Modern Superabsorbent Polymer Technology (John Wiley & Sons, 1997). Jensen, O. M. & Hansen, P. F. Water-entrained cement-based materials: I. Principles and theoretical background. Cement and Concrete Research 31, 647-654, (2001). Jensen, O. M. & Hansen, P. F. Water-entrained cement-based materials: II. Experimental observations. Cement and Concrete Research 32, 973-978, (2002). Krause-Rehberg, R. & Leipner, H. S. Positron Annihilation in Semiconductors. Vol. 127 (Springer, 1999). Charlton, M. & Humberston, J. W. Positron Physics. (Cambridge University Press, 2001). Mogensen, O. E. Spur reaction model of positronium formation. The Journal of Chemical Physics 60 (1974). Tao, S. J. Positronium Annihilation in Molecular Substances. The Journal of Chemical Physics 56 (1972).


14 Materials Characterisation V [15] Eldrup, M., Lightbody, D. & Sherwood, J. N. The temperature dependence of positron lifetimes in solid pivalic acid. Chemical Physics 63, 51-58, (1981). [16] Gidley, D. W. et al. Positronium annihilation in mesoporous thin films. Physical Review B 60, R5157 (1999). [17] Gidley, D. W. et al. Determination of pore-size distribution in lowdielectric thin films. Applied Physics Letters 76 (2000). [18] Pascual-Izarra, C. et al. Advanced fitting algorithms for analysing positron annihilation lifetime spectra. Nuclear Instruments and Methods in Physics Research Section A: Accelerators, Spectrometers, Detectors and Associated Equipment 603, 456-466, (2009). [19] Patro, A. P. & Sen, P. Anomalous parapositronium lifetime in water at 20 degrees C. Journal of Physics C: Solid State Physics 5, 3273 (1972). [20] Consolati, G. & Quasso, F. Evolution of porosity in a Portland cement paste studied through positron annihilation lifetime spectroscopy. Radiation Physics and Chemistry 68, 519-521 (2003). [21] Consolati, G. & Quasso, F. A positron annihilation study on the hydration of cement pastes. Materials Chemistry and Physics 101, 264-268 (2007). [22] Consolati, G., Dotelli, G. & Quasso, F. Positron lifetime spectroscopy as a probe of nanoporosity of cement-based materials. Radiation Physics and Chemistry 58, 727-731 (2000). [23] Salgueiro, W., Somoza, A., Cabrera, O. & Consolati, G. Porosity study on free mineral addition cement paste. Cement and Concrete Research 34, 9197, (2004). [24] Myllylä, R. & Karras, M. Positron Annihilation Probing for the Hydratation Rate of Cement Paste. Applied Physics 7, 303-306 (1975). [25] Sen Wang, P. U. et al. 1H nuclear magnetic resonance characterization of Portland cement: molecular diffusion of water studied by spin relaxation and relaxation time-weighted imaging. Journal of Material Science 33, 3065-3071 (1998).



15

Multi-technique investigation of calcium hydroxide crystals at the concrete surface E. Gueit1, E. Darque-Ceretti1, P. Tintillier2 & M. Horgnies2 1

MINES ParisTech, Center for Material Forming, Sophia Antipolis, France 2 Lafarge Centre de Recherche, St Quentin Fallavier, France

Abstract The durability and aesthetic qualities of high-performance concrete, which makes it particularly suitable for architectural applications, are constantly compromised by environmental aggressions. In this study, an innovative solution was developed to protect the concrete from these aggressions, which consists of growing a mineral coating on the concrete surface. The coating is composed of layered calcium hydroxide crystals, whose nucleation and growth are triggered byvarious non-ionic surfactants (the details of the process will not be presented). This paper describes the procedure used to investigate the structure of the formed crystals. Scanning Electron Microscopy and optical microtopography were used to determine the morphology of the crystals. Image analysis allowed the quantification of their amount, size and shape. The contribution and limits of each technique are discussed. Keywords: concrete, scanning electron microscopy, image analysis, optical microtopography, surface.

1 Introduction The mechanical and aesthetic durability of concrete is often compromised by the constant environmental aggressions to which the structures are exposed (organic or inorganic particles, algae, micro-organisms, staining from various sources). It is possible to protect concrete from these attacks and increase its durability by applying organic coatings on the hardened surface, but this comes with operational and environmental costs.

WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line) doi:10.2495/MC110021

16 Materials Characterisation V An innovative solution was proposed [1], which consists on covering and protecting high-performance concrete (HPC) with a mineral coating made of calcium hydroxide crystals (CH). The crystals growth happens during the concrete setting and is triggered by the presence of non-ionic surfactants at the concrete/formwork interface. One of the difficulties of this study was to determine the adequate characterization methods to properly investigate the influence of various surfactants on the amount, morphology and size of the crystals. Several techniques exist to observe, measure and study concrete hydrates, but not all of them are directly suitable for surface investigation. An original procedure had to be developed, combining several techniques. The structure and morphology of the crystals were assessed through Scanning Electron Microscopy (SEM) both on the concrete surface and on polished section. The SEM observations were completed by microtopography on the concrete surface. The quantification of the amount and size of the crystals was made by image analysis on binocular images of the concrete surface. The purpose of this paper is to illustrate what each of these techniques can bring to concrete surface studies.

2 Morphology of the crystals The most common method to observe concrete hydrates morphology is Scanning Electron Microscopy [2]. Most of the time, the observations are made on polished section in secondary electron mode. When higher magnifications or no polishing are required, fresh fractures can also be observed. Surprisingly, it is very rare to find published picture of concrete surface directly observed by SEM. In this study, SEM observations were conducted on both surfaces and crosssections of the concrete. Small cubes (1 cm x 1 cm x 1 cm) were cut from each concrete sample. For surface observation, the cubes were directly carbon-coated on the adequate face and observed in secondary electron mode. For crosssections observations, the cubes were impregnated, polished and carbon-coated, and then observed in back-scattered electron mode. A SEM FEG Quanta 400 from FEI Company was used at an accelerating voltage of 15kV and current intensity of 1mA.

Figure 1:

SEM observations of a standard high-performance concrete surface in secondary electron mode (A) and in back-scattered electron mode on a polished section (B).



17

Figure 1 shows observations of both the surface (A) and the cross-section (B) of a standard high-performance concrete. No specific features are visible on the surface, except for a couple of scratch due to the formwork defects. As for the polished section, it reveals a classical concrete microstructure, with dark grey aggregates and light grey un-hydrated cement grains surrounded by cement paste. Figure 2 shows SEM observation of two cross-sections from two different samples. These observations reveal the presence at the concrete surface of a thin layered structure presenting various orientations and organization, sometimes well-aligned parallel to the surface (A), sometimes arranged in a more chaotic way in the first micrometers of the surface (B). At this level of observation, it is not possible to clearly indentify the nature of this unexpected phase. The limited magnification in back-scattered mode, the thin structure of the hydrates and the surface damages due to the polishing all complicate the interpretation. This is why it is necessary to make complementary observations of the surface itself.

Figure 2:

SEM observation in back-scattered electron mode of two polished sections from two different concrete samples from the study.

Figure 3 shows the same samples as figure 2 observed from the surface in secondary electron mode. These micrographs allow a better understanding of the structure of the crystals. All of them share the same layered and flaky structure with different orientations. In all cases, the crystals are composed of thin leaves that grow around the nucleation point. For some crystals, these leaves are strictly parallel to the surface and grow as ‘flower petals’ around the central point, reaching sometimes a perfect hexagonal shape. For other crystals, the leaves are strictly perpendicular to the surface, forming a very regular spherulitic structure. This very organized structure was not visible on the polished sections, where the leaves appeared randomly implanted in the surface. This is due to the polishing, which damaged the first micrometers of the surface and disturbed the structure. The characteristic hexagonal shape of the crystals, as well as their layered organization, allow identifying them as calcium hydroxide – no other concrete phase would present this morphology. This is confirmed by EDS analysis, where mostly calcium is detected is these areas (spectrum not showed). WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


Figure 3:

SEM observation in secondary electron mode of the surface of two different samples from the study.

3 Amount and size of the crystals 3.1 Microtopography SEM observations were very useful to highlight the presence of an unusual phase and to identify it as CH crystals, but they are not sufficient to get a complete insight on the structure of the crystals. Even though the secondary electrons give information on the surface topography, it is difficult to evaluate from the micrographs how thick the crystals are or how deep they are embedded in the surface. As for the polished section, the crystals layers are too strongly delaminated by the polishing for the observation to be conclusive. This is why the observations were completed with profilometry of the surface. Mechanical and optical profilometry are sometimes used on concrete to assess its roughness and evaluate its behavior regarding adhesion problems [3–5]. In our case, the measurements were made using optical profilometry only, because a mechanical probe is likely to damage the very fragile CH crystals. The measurements were made on a confocal full-field 3D surface profilometer with a spot of 2 micrometers and a working distance of 4.5 mm. In these conditions, the vertical resolution is 0.01 micrometers and the lateral resolution 0.1 micrometer. Areas of 4 x 4 mm² were scanned with a step of 10micrometer. The data were computed using the software MountainsMap. In the examples given below, there were only two steps of data treatments: the maps were straightened to compensate for the horizontality defects, and profiles were extracted. Figure 4 shows an example of profilometry on a sample where the crystals are oriented preferentially parallel to the surface. Two patterns appear on the 2D mapping: regular and slightly curved lines, which are due the formwork texture, and irregular spots, which are the CH crystals. Below the mapping is a roughness profile extracted along the dotted line. The narrow peaks correspond to the lines, the larger one at 1.7 mm corresponds to the CH crystal in the middle. The profile allows measuring the size of the crystal (400 micrometers) and the height between the concrete surface and the top of crystal (5 micrometers). WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


19

Figure 4:

Example of a mapping of the concrete surface. The profile was extracted from the mapping along the black dotted line.

Figure 5:

Example of a mapping of the concrete surface. The profile was extracted from the mapping along the black dotted line.


20 Materials Characterisation V The same measurements can be made of figure 5. The mapping was acquired on the same sample, but in an area were the CH crystals have been pulled out of the surface during the demoulding phase, leaving hollows the size of the crystals. The depth of the hollows gives the height between the surface and the bottom of the crystals. The measurements from figures 4 and 5 give an estimation for the total thickness of the crystals of 20 micrometers. Of course, the operation has to be repeated a large number of times to obtain statistical results. This is only one example taken from one of the samples. It should be noted that profilometry gives no conclusive information for the samples where the crystals are oriented perpendicular to the surface. Figure 6 shows a mapping of such a sample (A): the spherulitic structures appear slightly lighter than the surface, but are difficult to distinguish. A zoom on the mapping (B) reveals that the crystals are indeed apparent, but that the resolution is far from being sufficient to separate the different flakes composing the structure.

Figure 6:

Example of a mapping from a sample with perpendicular crystals. The 3D mapping on the right is a zoom from the 2D mapping on the left.

3.2 Image analysis The only way to know the amount of CH crystals as well as their dimension and geometrical parameters is to count and measure them one by one. Image analysis software are capable of automatically achieve this procedure on a large number of images, provided that they are correctly settled by the operator. Image analysis has been successfully used in previous studies [2, 6] to assess the amount of calcium hydroxide in concrete as well as the geometrical parameters of the particles (size, shape, etc.). In those cases, image analysis was conducted on SEM micrographs from polished sections. In our case, there is no need to work on sections as the crystals are on the concrete surface. Furthermore, the crystals are so large that SEM micrographs are not suitable for a proper counting. Even at low magnification, there are too few particles in the observation field. On the contrary, images taken under a binocular have the right scale to count and measure the particles. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


21

The CH crystals are shinier than the cement paste, which makes them visible with the naked eye, but this is not sufficient for image analysis software to distinguish them. To overcome this problem, the samples were coloured with a black felt-tip pen prior to analysis. The black ink penetrates the cement paste, strongly colouring it, but is not absorbed by the crystals which then become clearly distinguishable. For each concrete samples, five small areas of the surface (1 cm2) were coloured and photographed under a binocular. The images were treated and analysed using the open-source software ImageJ [7]. The most difficult step in image analysis is to separate the studied particles from the background. In our case, the following procedure was used. Figure 7 shows an example of an image prior to treatment. The brightness and contrast of the images were optimized using the automatic function on the software. The optimization was not based on the whole image, but on the histogram analysis of a small area of one of the particle on the image. This step created a highly contrasted image where particles appeared red and the background black (figure 8).

Figure 7:

Figure 8:

Example of an image before treatment by the software imageJ.

Example of an image after optimization of the contrast and brightness.


22 Materials Characterisation V The noise was removed using the ‘remove outliers’ function of the software, which replaces each pixel by the median value of its neighbours in a given area (figure 9). In this case, a radius of 20 pixels was considered, and the removing was set to occur if the difference between the considered pixel and the median of its neighbour was higher than 10.

Figure 9:

Example of an image after noise removal.

The images were then converted to binary and inverted. At this point, the background appeared white and the particles black (figure 10).

Figure 10:

Example of an image after conversion to binary.

The holes in the particles were closed using the ‘Fill holes’ function of the software (figure 11). Finally, the ‘watershed’ function was used to automatically draw the outlines of the particles (figure 12). This step might be the main cause of errors in the measurements, as watershed segmentation works best for convex objects that do not overlap too much. In some cases, the software was not able to correctly separate clusters of particles, which were then counted as one big particle.



Figure 11:

Figure 12:

23

Example of an image after filling the holes.

Example of an image after the software automatically delimited the particles.

Once the image was treated, the particles were counted one by one. The smallest ones (with a surface smaller than 0.1 mm²) were not considered as they are more likely residual noise than actual crystals. The particles overlapping with the boarder of the image were not considered either as their size would not be correctly measured. This means than the measured amount of crystals (from 1 to 20 particles per mm² depending on the samples) is slightly smaller than reality. The following parameters were measured for each particle: - The coordinates of the centre of the particles. These coordinates were used to calculate the distance between each crystal and its closest neighbour and to verify that the nucleation appeared randomly and homogeneously on the surface (figure 13). - The Feret’s diameter, which is the longest distance between two opposing points of the particle. - The fraction area of the surface covered by the crystals. As these crystals are artificially grown to act as a mineral coating, this fraction area must be as high as possible. The results showed that even if the crystals seemed very large, they hardly cover 50% of the surface, except in one promising case where almost 100% of the surface was coated. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)

24 Materials Characterisation V -

The circularity of the particle, which is defined by 4.area / perimeter². A perfect circle has a circularity of 1, whereas a very elongated particle has a circularity close to 0 (figure 14). This geometrical parameter is very convenient to distinguish the crystals that grow parallel to the surface with regular shapes (high circularity) from the crystals that grow perpendicular to the surface, which have a low circularity.

Figure 13:

Example of spatial distribution of the particles on the concrete surface, showing that the nucleation occurs rather homogeneously on the surface.

Figure 14:

Two examples of the circularity distribution. The curve marked with squares corresponds to a sample where the crystals have a low circularity because they grow perpendicular to the surface. The curve marked with dots corresponds to a sample where the crystals have a high circularity because they grow parallel to the surface and adopt very regular shapes.



25

The results of the image analysis allowed a better understanding of the influence of various surfactants on the nucleation (amount of particles) and growth (size and orientation of the particles) of the CH crystals. Two types of sample could not be properly analyzed though, one because it exhibited excessive roughness that induced too much background noise on the images, and one because the crystals overlapped too much to be clearly separated. In those cases, a more sophisticated procedure has to be developed. In particular, software capable of separating overlapping particles exists, but they were not tested yet in this study.

4 Conclusion Three techniques – scanning electron microscopy, microtopography and image analysis – were successfully used to assess the morphology and geometrical parameters of calcium hydroxide crystals at the concrete surface. These techniques are well-known and developed, but not necessarily widely used in concrete research. Yet, they have proved very efficient in this case to make a preliminary study of a new phenomenon – the massive growth of CH crystals in presence of surfactants, bringing complementary information and results. Of course, they absolutely do not make further investigation any less necessary. For example, a proper crystallographic study would be essential to fully understand the growing mechanisms and the action of the surfactants. In the field of materials characterization, it is important to be creative and to combine and adapt existing techniques. This is particularly true in the field of concrete research, and even more when it comes to concrete surface, a topic which is slowly emerging and where a lot of fascinating research still waits to be done.

References [1] Gueit, E., Darque-Ceretti, E., Tintillier, P. & Horgnies, M., Surfactantinduced growth of calcium hydroxide at the concrete/formwork interface as a mineral coating for concrete, Manuscript submitted for publication. [2] Skalny, J., Gebauer, I. & Odler, I., (eds). Calcium Hydroxide in Concrete, The American Ceramic Society: Westerville, 2001. [3] Garbacz, A., Courard, L., & Kostana, K., Characterization of concrete surface roughness and its relation to adhesion in repair systems, Materials Characterization, 56, pp. 281-289, 2006. [4] PER09 Perez, F., Bissonette, B. & Courard, L., Combination of mechanical and optical profilometry techniques for concrete surface roughness characterisation. Magazine of Concrete Research, 61(6), pp. 389-400, 2009. [5] Ramirez, A.M., Demeestere, K., De Belie, N., Mäntylä, T., & Levänen, E, Titanium dioxide coated cementitious materials for air purifying purposes: Preparation, characterization and toluene removal potential, Building and Environment, 45, pp. 832-838, 2010.


26 Materials Characterisation V [6] Gallucci, E. & Scrivener, K., Crystallisation of calcium hydroxide in early model and ordinary cementitious systems, Cement andConcrete Reseach,37, pp. 492-501, 2007. [7] NIH, http:\\rsbweb.nih.gov/ij



27

Characterization of the influence of the casting mould on the surface properties of concrete and on the adhesion of a protective coating M. Horgnies, P. Willieme, O. Gabet, S. Lombard & M. Dykman Lafarge Centre de Recherche, St Quentin-Fallavier, France

Abstract Protective coatings are deposited on concrete to improve aesthetics and to prevent ageing. However, their adhesion on concrete depends on several interlinked parameters. In this study, the surfaces of concrete are characterized according to the process of casting and post-treatment used (sandblasting) by using Scanning Electron Microscopy (SEM), Fourier Transformed-Infrared (FTIR) spectroscopy and profilometry. The surface properties are correlated to the adhesion force of a polyurea (PU) coating. The development of a specific peel test (a strengthened and porous membrane is introduced into the layer of liquid coating before its crosslinking) ensures a reproducible debonding of the coating/concrete system and allows measuring the fracture energy. Moreover, the interface after debonding is analyzed by FT-IR to highlight the presence of concrete/coating residues and to determine the locus of failure. Results underline that the nature of casting mould influences the concrete surface and modifies the adhesion of PU coating. The mould made of polyoxymethylene (POM) induces a micro-tearing of the extreme surface of concrete during demoulding. By increasing the roughness and the open porosity of the concrete surface, this tearing enhances the adhesion of the coating. On the contrary, the smooth concrete surface, induced by the use of a polyvinylchloride (PVC) mould, decreases the anchorage of the coating. Finally, the sandblasting of the surface, by increasing the roughness and the interface area, is an interesting treatment to promote the adhesion of PU coating, whatever the mould used for the casting. Keywords: concrete, coating, roughness, FT-IR, SEM, peel test, adhesion.



1 Introduction The staining of concrete could occur due to its specific microstructure that retains the liquid and dust particles. The deposition of a coating is then important to close the surface porosity and protect concrete against acid rains, settlement of algae and lichens…etc [1, 2]. PU coatings are commonly used due to their high resistance against chemical and mechanical aggressions [3, 4]. However, the surface properties of concrete depend on several interlinked parameters as chemical composition, intrinsic porosity and roughness. The aim of this study concerns the influence of the casting process and post-treatment on the adhesion between PU coating and concrete surface. The influences of the casting conditions on the hardened concrete surface were characterized by SEM, FT-IR and profilometry. Secondly, this study was undertaken to determine if the surface properties of concrete could influence the adhesion of PU coating. Fracture energies were measured by a specific 90°-peel test. This method was retained because it is appropriate to characterize the adhesion of thin films [5–8]. Some publications have already showed the use of strengthened membrane or mesh sheet that were incorporated into the bulk of soft material to characterize [9]. Concerning our system, the introduction of a polymer membrane into the bulk of the coating was necessary to strengthen the system and measure a reproducible adhesion of the PU coating. After the peel tests, the FT-IR analyses of the debonded faces were undertaken to detect the residues of concrete or coating and determine the loci of failure. FT-IR spectroscopy allows detecting organic compounds of coatings [10, 11] and several components of concrete [12].

2 Material and methods 2.1 Material 2.1.1 Substrates made of hardened concrete A high-performance concrete was prepared by mixing 31% of white Portland cement (CEM I 52.5 PMES from Lafarge), 9% of limestone filler (DURCAL 1), 7% of silica fumes (MST), 43.5% of sand (BE01) and 1.5% of admixture. A water to cement ratio (W/C) of 0.26 was used. The samples were prepared by pouring the fresh concrete mixture into horizontal and rectangular formwork (15x12x1 cm) made of PVC or POM. The concrete samples were removed from their formworks after 18 hours and were stored during 28 days under ambient conditions (25°C; 50% relative humidity) to complete their hydration. Some concrete samples were then sandblasted after demoulding (by using a powder of corindon) to increase their roughness. The sandblasted samples were cleaned by air flow to remove the dust before the deposition of coating. 2.1.2 Coating and conditions of deposition The PU coating was composed of 50% of isocyanates diluted into 45% of solvent (butyl acetate). A catalyst (dibutylétain laurate, DBTL) was added into the mix to initiate the reaction with water (present in concrete or in atmosphere). WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


29

The isocyanates units reacted with water to produce a polyurea-based film [13]. The coatings were sprayed (120 g/m²) with air pressure of 3 atmospheres on the concrete surface. The drying period was about 3 days under ambient conditions (25°C, 50% relative humidity) before peeling. 2.2 Methods of characterization 2.2.1 Scanning electron microscopy (SEM) Samples were characterized by using a high-resolution field-effect gun digital scanning electron microscope (SEM FEG Quanta 400 from FEI Company; using an accelerating voltage of 15 keV and a current intensity of 1 nA). Images of the cross-sections were obtained after being polished. 2.2.2 Profilometry The roughness of concrete samples was measured with a Surftest SJ-201 M mechanical profilometer (Mitutoyo) in order to calculate the arithmetic mean of the profile deviations from the mean line (Ra). The Ra value was obtained by compiling the arithmetic mean of 5 profiles of 12.5 mm. 2.2.3 Fourier transform-infrared spectroscopy (FT-IR) The FT-IR spectrometer Nicolet iS10 (Thermo Fisher Scientific Inc.) was equipped with a deuterated triglycine sulfate (DTGS) detector and controlled by OMNIC software. The Attenuated Total Reflexion (ATR) mode was mainly used in this study. FT-IR (in ATR mode) characterized the sample over a thickness of a few µm. The sampling area analyzed was approximately 1 mm². The crystal used was made of diamond and 16 scans were routinely recorded over the range 4,000-650 cm-1 with a spectral resolution of 4 cm-1. The background was collected at ambient atmosphere before analyzing each sample. Spectra were corrected with a linear baseline. No specific preparations of the samples of concrete and coating were performed before FT-IR analyses: they were studied just after demoulding or after debonding. 2.2.4 Specific peel tests of concrete/coating system Peel test allows measuring the debonding force (F). According to the peel angle (θ) and the width of the adhesive coating (w), the fracture energy (G) could be calculated according to [14]: F G w 

 1  cos    

In the specific case of 90° peel angle, G is equal to the peeling force (F) divided by the width (w) of the adhesive coating. The coating was strengthened by inserting a membrane. The membrane was firstly cut in polyamide 6.6 tissues (Nytex from Dutscher SAS) and deposited on each concrete samples. The coating was directly sprayed on this membrane in order to avoid any air bubble at the interface. Finally, the edges of the membrane were cut to avoid any shear during the debonding. The membrane was flexible but non-stretched under the


30 Materials Characterisation V solicitations that occurred during the peel test. The dimensions of each membrane were 200 mm long, 20 mm wide, 120 µm thick, with an open porosity of 50% and a mesh opening of 150 µm. We observed by SEM (Figure 1) a crosssection of the concrete/coating interface in order to describe the system. No mechanical step was used to initiate the crack at the interface. All the 90°-peel tests were performed with a 1000 N sensor, equipped with a specific mobile table (as detailed by Figure 2). All the peel tests were performed under ambient conditions (25°C, 50% relative humidity) and by using a constant speed of peel (0.2 mm/s). Strengthened membrane

PU coating (130 µm thickness)

Concrete surface

Figure 1:

Figure 2:

SEM image of a cross-section of the concrete/coating system.

Schema and image of the experimental test of peeling.



31

3 Results 3.1 Analyses of the surfaces of reference materials 3.1.1 Concrete surfaces according to the nature of the mould Table 1 details the IR bands of concrete (demoulded with PVC). Portlandite (at 3640 cm-1) and C-S-H*/silica (large band at 1078 cm-1) were well detected. Moreover, three strong bands were also assigned to the presence of CaCO3. Their detection confirmed previous studies [12]. Detection of CaCO3 could be due to the carbonation process but also due to the use of limestone filler in concrete mix. As presented by Table 1, concrete demoulded with POM (spectrum not shown) underlines distinct IR bands at 935 and 900 cm-1 that were assigned to alite [15] (a component of the cement before hydration). This FT-IR spectrum showed also peaks at 1092 and 800 cm-1, assigned to Si-O bonds from silica fumes [12]. These silica fumes and alite are usually present into the bulk of concrete sample. Moreover, no IR band of the portlandite (Ca(OH)2) was detected. Table 1:

Assignment of FT-IR bands recorded on reference materials. FT-IR bands (cm-1)

Concrete (PVC mould)

O-H; Ca(OH)2

3640

X

C=O; CaCO3

1410; 872; 710

X

X

Si-O; silicates, CSH

1080-970

X

X

X

X

Assignment

Concrete (POM mould)

PU coating

Si-O; C3S

935; 900

Si-O; silica

797; 777

X

CH2/CH3; methyl units

2850; 2950; 1450

X

N-H; urea

3335

X

NCO; isocyanates

2270

X

C=O; urea

1690

X

The concrete samples were also studied by SEM. Concrete demoulded with POM (Figures 3a and 3b) appears to be rougher than concrete demoulded with PVC (Figures 3c and 3d). The flat topography observed after using PVC mould contrasted with the heterogeneous surface obtained after using a POM mould. These observations confirmed the measurements obtained by profilometry (Table 2): the Ra values of concrete could vary according to the demoulding process. Indeed, the results obtained by SEM suggest that a tearing of the extreme surface of concrete could happen during the removing of the POM mould. This hypothesis could explain why silica fumes were easily observed by SEM (Figure 3b) and why alite was detected by FT-IR spectroscopy (Table 1).


32 Materials Characterisation V (a)

(b)

Silica fumes

(d)

(c)

Figure 3:

SEM images recorded on concrete samples: (a, b) if POM mould; (c, d) if PVC mould.

Table 2:

Roughness of concrete according to the demoulding and posttreatment.

Ra roughness (µm)

Concrete (POM mould)

Reference

0.7 µm +/- 0.3

Concrete (PVC mould) 0.5 µm +/- 0.3

After sandblasting

12.2 µm +/- 0.4

12.1 µm +/- 0.5

3.1.2 Concrete surfaces after sandblasting Table 2 summarizes the roughness of sandblasted concrete, compared to the one before demoulding. The initial roughness of 0.5-0.7 µm range (depending on the nature of mould used) increased to a higher Ra values of 12 µm after sandblasting (whatever the nature of the mould used). SEM images of the sandblasted concrete, presented in Figure 4, could be compared to the ones of the reference concrete samples (Figure 3). The sandblasted surface was rougher than after demoulding. Silica fumes were also detected after sandblasting. Figure 5 compares FT-IR spectra recorded on concrete (PVC mould) samples before and after sandblasting. The intensity of the bands assigned to the CaCO3 compounds (at 1410, 870 and 710 cm-1) decreased after sandblasting while the intensity of peaks of alite (at 930-900 cm-1) and silica (at 1100 cm-1) increased. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


33

Silica fumes

Figure 4:

SEM images of concrete (PVC mould) samples after sandblasting.

CaCO3

(b)

(a) Silica

Figure 5:

Alite

FT-IR spectra of concrete samples (PVC mould): (a) before sandblasting; (b) after sandblasting.

3.1.3 PU coating with membrane FT-IR spectroscopy is known to determine the degree of crosslinking of the isocyanates because a band at 2270 cm-1 characterizes presence of isocyanates units [16, 17]. In this study, no results of adhesion were linked to the degree of crosslinking of the coating. However, Table 1 underlines other IR bands assigned to the presence of urea units (at 3335 and 1690 cm-1). These bands were used to detect the residues of PU coating after peeling. The membrane made of WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)

34 Materials Characterisation V polyamide 6.6. was also analyzed by FT-IR in order to assign the reference peaks (amide units at 3300, 3070, 1630 and 1530 cm-1). However, these IR bands of the polyamide were never detected at the interface after debonding. 3.2 Results of peel tests and FT-IR analyses of the interface after debonding 3.2.1 Influence of the mould used to cast the concrete Figure 6 compares the peel curves obtained during the debonding of PU coating from concrete samples. The adhesion of the coating varied according to the nature of the casting mould. 1400

1200

(a) Concrete (POM mould)

G (N/m)

1000

800

600

(b) Concrete (PVC mould)

400

200

0 0

10

20

30

40

50

60

70

80

Length of debonding (mm)

Figure 6:

Peel curves of coating recorded on concrete samples: (a, in blue) if POM mould; (b, in red) if PVC mould.

As described by Table 3, the fracture energy measured during the peel of the PU coating from concrete (POM mould) is higher (1000 N/m +/- 150) than the one measured during the peel from concrete (PVC mould): 400 N/m +/-50. Figure 7a presents the FT-IR spectrum recorded on concrete (POM mould) side after peeling: almost all the IR bands were assigned to the residues of coating that covered the entire surface. These results confirm that the high fracture energy was linked to a debonding into the bulk of the coating. It could also be deduced that the adhesion between concrete and coating was better than 1000 N/m. On the contrary, Figure 7b shows FT-IR spectrum of concrete (PVC mould) side after peeling: compounds of concrete (such as CaCO3, silica and portlandite) were mainly detected. Concerning the internal side of coating after peeling from concrete (PVC mould) sample, all the IR bands (spectrum not shown) were assigned to the own components of the coating. No FT-IR bands could be assigned to concrete residues.



Table 3:

Sample

35

Results obtained by peel tests and by FTIR analyses of the interface. Post-treatment of concrete before coating

Fracture energy (N/m)

Failure localization deduced from FTIR analyses

/

1000 +/- 150

Inside the PU coating

Sandblasting

1200 +/- 100


/

400 +/- 50

At concrete/PU interface

Sandblasting

1200 +/- 100


Concrete/POM

Concrete/PVC

(a) Concrete side (POM mould) after debonding

(b) Concrete side (PVC mould) after debonding

Figure 7:

FTIR spectra of concrete sides after peeling: (a) if POM mould; (b) if PVC mould.

These data allow concluding that the debonding occurred at the interface between concrete and coating. The fracture energies and FT-IR analyses of the locus of failure could be correlated to the SEM observations of concrete surface WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)

36 Materials Characterisation V before the deposition of the coating. Figure 3 had highlighted that the surface of concrete demoulded with POM was more heterogeneous than concrete demoulded with PVC, which was smoother. We can deduce that the liquid coating (before crosslinking) strongly diffused and anchored into the heterogeneous concrete surface, demoulded with POM. On the contrary, the smooth and close concrete demoulded with PVC did not allow a high mechanical anchorage of the coating. 3.2.2 Influence of the sandblasting of concrete surface The peels tests undertaken on the sandblasted concrete (PVC mould) highlighted high fracture energies (about 1200 N/m) while the peel undertaken on reference concrete (PVC mould) samples induced a low level of adhesion of the PU coating (350-400 N/m), as described in Figure 8a. Concerning concrete demoulded with POM, the fracture energies (presented in Figure 8b) did not 1600

(a)

1400

Sandblasted concrete (PVC mould)

1200

G (N/m)

1000 800 600

Not-sandblasted concrete (PVC mould)

400 200 0 0

10

20

30

40

50

60

70

80

Lenght of debonding (mm)

1600

(b) 1400

Sandblasted concrete (POM mould)

1200

G (N/m)

1000

Not-sandblasted concrete (POM mould)

800 600 400 200 0 0

10

20

30

40

50

60

70

80

Length of debonding (mm)

Figure 8:

Peel curves of the coating recorded on concrete with or without sandblasting: (a) if PVC mould; (b) if POM mould.



37

evolve: they were as high after sandblasting as after demoulding. Thanks to these results, we can conclude that a sandblasting of concrete demoulded with PVC could be interesting to enhance the adhesion of PU coating. As described by SEM images (Figure 4) and values of roughness (Table 2), the open porosity and the roughness of concrete demoulded with PVC were higher after the sandblasting (Figure 4c) than after demoulding (Figure 3c). On the contrary, the sandblasting was less helpful with concrete demoulded with POM because the open porosity was already present after demoulding. FT-IR analyses of concrete sides were performed after peeling but only residues of coating were detected on these concrete sides (whatever the nature of mould used). The debonding occurred inside the layer of the PU coating if the substrate had been sandblasted before the deposition of the coating.

4 Conclusion A specific methodology of 90°-peel test was used to measure the fracture energy between concrete and PU coating. By introducing a porous and thin membrane to strengthen the layer of coating, measurements were reproducible and allowed comparing the fracture energies with the FT-IR analyses of the loci of failure recorded after debonding. These results were correlated to the surface properties of the reference concrete surface after demoulding (topography, composition, roughness) that were characterized by several methods. The influence of the nature of mould was significant. Some moulds, such as POM, could induce a micro-tearing of the extreme surface of concrete during the demoulding. This phenomenon increased the roughness (and the open porosity) of concrete and improved the anchorage of the liquid coating (before its crosslinking). On the contrary, the smooth and flat concrete surface induced after using a PVC mould did not allow a good adhesion of PU coating. The influence of the mechanical anchorage was confirmed by the high fracture energies measured on the sandblasted samples of concrete. In future, analyses of the surface porosity and sorptivity will be performed to confirm these results.

References [1] Dubosc, A., Escadeillas, G. & Blanc, P.J., Characterization of biolical stains on external concrete walls and influence of concrete ad underlying material. Cement and Concrete Research, 31, pp. 1613-1617, 2001. [2] Manoudis, P.N., Karapanagiotis, I., Tsakalof, A., Zuburtikutis, I., Kolinkeova, B. & Panayiotou, C., Surface properties of superhydrophobic coatings for stone protection. Journal of Nano Research, 8, pp. 23-33, 2009. [3] Chattopadhyay, D.K. & Raju, K.V.S.N., Structural engineering of polyurethane coatings for high performance applications. Progress in Polymer Science, 32, pp. 352-418, 2007.


38 Materials Characterisation V [4] Carmona-Quiroga, P.M., Martínez-Ramírez, S., Sobrados, I. & BlancoVarela, M.T., Interaction between two anti-graffiti treatments and cement mortar (paste). Cement and Concrete Research, 40, pp. 723-740, 2010. [5] Barquins, M. & Ciccotti, M., On the kinetics of peeling of an adhesive tape under a constant imposed load. International Journal Adhesion and Adhesives, 17, pp. 65-68, 1997. [6] Horgnies, M., Darque-Ceretti, E. & Combarieu, R., Adhesion of pressure sensitive adhesives to automotive coatings: Influence of topcoat composition. Journal of Adhesion Science and Technology, 18, pp. 10471061, 2004. [7] Johnson, K.L., Kendall, K. & Roberts, A.D., Surface Energy and the contact of elastic solids. Proceedings of Royal Society of London, 324, pp. 301-313, 1971. [8] Kinloch, A.J., Adhesion and adhesives: science and technology. Chapman and Hall, London, New York, pp. 66-73, 1987. [9] Giannis, S., Adams, R.D., Clark, L.J. & Taylor, M.A., The use of a modified peel specimen to assess the peel resistance of aircraft fuel tank sealants. International Journal of Adhesion and Adhesives, 28, pp. 158-175, 2008. [10] Almeida, E., Balmayore, M. & Santos, T., Some relevant aspects of the use of FT-IR associated techniques in the study of surfaces and coatings. Progress in Organic Coatings, 44, pp. 233-242, 2002. [11] Poliskie, M. & Clevenger, J.O., Fourier Transform Infrared (FT-IR) spectroscopy for coating characterization and failure analysis. Organic Finishing, pp. 44-47, 2008. [12] Chollet, M., Horgnies, M., Analyses of the surfaces of concrete by Raman and FT-IR spectroscopies: comparative study of hardened samples after demoulding and after organic post-treatment. Surface and Interface Analysis, In press. [13] Agrawal, R.K. & Drzal, L.T., Adhesion mechanisms of polyurethanes to glass surfaces. Part I. Structure property relationships in polyurethanes and their effects on adhesion to glass. Journal of Adhesion, 54, pp. 79-102, 1995. [14] Kendall, K., Thin-film peeling – the elastic term. Journal of Physics D: Applied Physics, 8, pp. 1449-1452, 1975. [15] Frost, R.L., Cejka, J. & Weier, M.L., Molecular structure of the uranyl silicates – a Raman spectroscopic study. Journal of Raman spectroscopy, 37, pp. 538-551, 2005. [16] Daniel-da-Silva, A.L., Bordado, J.C.M. & Martin-Martinez, J.M., Moisture curing kinetics of isocyanate ended urethane quasi-prepolymers monitored by IR spectroscopy and DSC. Journal of Applied Polymer Science, 107, pp. 700-709, 2007. [17] Agrawal, K. & Drzal, L., Adhesion mechanisms of polyurethanes to glass surfaces. Journal of Adhesion Sciences Technology, 9, pp. 1381-1400, 1995.


Section 2 Nano-materials



41

HRTEM techniques applied to nanocrystal modeling: towards an “atom-by-atom” description D. G. Stroppa1,2, L. A. Montoro1, E. R. Leite3 & A. J. Ramirez1,2* 1

Brazilian Synchrotron Light Laboratory, Brazil Mechanical Engineering School, University of Campinas, Brazil 3 Department of Chemistry, Federal University of São Carlos, Brazil 2

Abstract The development of technologies based on nanostructures presents a wide range of challenges for materials scientists and engineers, including the attainment of well-controlled synthesis procedures, the improvement of characterization techniques down to the atomic scale resolution, and the conception and validation of reliable models that can describe materials properties as functions of their morphology and fabrication process. A relevant topic in this scenario is the correlation among the spatial distribution of chemical elements, the surface energy configuration, the growth mechanism, and the resultant nanocrystal 3D morphology. This work presents an overview on the use of advanced HRTEM techniques for the quantitative analysis of nanocrystals and how these results can be used to implement nanocrystals models, which can analytically describe the material features on an atomic level. The presented findings show the combined use of experimental data and theoretical tools, such as image simulation and ab initio surface energy calculations, for the advanced quantitative characterization of nanocrystalline systems. The combination of experimental and theoretical efforts on HRTEM characterization represents a powerful tool for the nanocrystal 3D morphology elucidation with atomic resolution and the chemical/structural properties assessment in a quantitative way. Thereby, it is presented as the stepping stone towards the development of novel approaches to describe nanostructured systems. Keywords: HRTEM, nanocrystals modeling, quantitative analysis. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line) doi:10.2495/MC110041


1 Introduction Nanotechnology has been considered one of the most promising branches on the scientific development since the 60’s [1] and its contributions to various fields are notorious nowadays. The “nano” suffix and its effects became an intrinsic subject of materials science and technology on the last few years as the unique properties from nanometer scale are becoming increasingly relevant in those fields. The traditional materials engineering perspective [2], depicted in Figure 1, can also be applied for the analysis of nanostructured systems. This approach considers that the design of materials aiming an optimum performance requires the complete understanding of the correlation among the processing routes parameters, the materials structure and the resultant properties. However, regarding nanostructured systems, this correlation may be stated as the interrelationship among the synthesis parameters, the atomic scale morphology and the energy configuration [3]. As the surface area to volume ratio is enhanced for such systems, unique properties related to the surface energy distribution become relevant. These properties are the source for the unique performance of nanomaterials [4].

Figure 1:

Engineering approach for describing and correlating materials features.

Materials characterization plays a crucial role on the development of reliable models for the description and the design of nanostructured materials with specific features for countless technological areas. Apart from the specific questions that different characterization techniques may answer, there are three general issues that can be highlighted as the most relevant questions for the nanostructured materials modeling and design. The first one is related to the morphology characterization at the atomic scale. As mentioned before, it is vital to assess the nanomaterials structure with high resolution in order to evaluate its correlation with the system properties. The ultimate goal in this sense is the “atom-by-atom” description of nanocrystals, WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


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which would be specially required for the determination of atomic species segregation on alloyed/doped materials [5] and the high resolution strain state analysis of nanostructures [6]. The second issue is related to the nanocrystals surface structure and its dependency on the synthesis environment and chemical species. This aspect is intrinsically correlated to the nanocrystals morphology and, consequently, to the material surface energy distribution and properties [7]. The third relevant point is the interrelation between the nanocrystals and the system growth behavior [8]. These aspects represent crucial issues for the understanding and optimization of synthesis methodology as they are closely related to the nanostructured system stability and long term reliability [9]. As describing the properties and behavior of each nanocrystal in an actual nanostructured system represents an overwhelming many-bodies problem, simplified models are needed for the engineering of nanomaterials in a reliable manner. Nanocrystal modeling [10] is a wide approach where nanostructured systems components are described as building blocks with specific morphology, energy distribution, and interaction mechanisms. In these terms, the development of accurate models may be a fundamental tool for nanocrystals and mesocrystals shapes prediction, design and control of growth processes, and the resulting properties tuning of an unlimited number of systems. The development of reliable modeling methodologies represents a hard task for materials scientists and engineers due to involved challenges, especially when quantitative high resolution characterization analyses are required. Even though advanced characterization techniques based on electron microscopy, scanning probe microscopy and synchrotron radiation are feasible and complimentary in this scenario, limitations [11] associated to each one of them prevent their isolated use a for reliable analysis. The most relevant limitations of high resolution techniques are summarized on Table 1. Table 1:

Most relevant limitations from high resolution characterization tools.

ISSUE Sample Preparation Sample Stability Energy Resolution Signal to Noise Ratio Sampling Averaging Data Analysis

EFFECT modification on the sample original features characteristics changes during analysis spectroscopic signature may have overlaps poor detection system sensitivity lack of statistical representativeness lack of features from individual particles deconvolution, modeling and fitting

Although continuous improvements on the characterization instrumentation and methodologies enhance the individual techniques performance, the combined use different available experimental techniques and theoretical approaches are widely used to overcome the inherent limitations of the quantitative analysis of nanosized systems. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)

44 Materials Characterisation V Transmission Electron Microscopy (TEM) based techniques stand out among other characterization techniques for nanostructured systems due to the possibility of a comprehensive and versatile approach including high resolution imaging, diffraction and spectroscopy. This multipurpose characteristic allied with simulation procedures provides complementary information that can be used to improve the analysis precision. Recent developments on the TEM instrumentation and analysis methods allowed outstanding advances on nanoscience and nanotechnology [12]. The spherical aberration (Cs) correction [13] can be considered among the most influential improvements on both high resolution TEM (HRTEM) and high resolution scanning TEM (HRSTEM) due to a number of factors. For HRTEM, the Cs-correction of the objective lens resulted in the drastic improvement on achievable spatial resolution, the possibility of larger gaps in the pole piece for in situ experiments, and new imaging modes [14] due to fine tuning of the aberration coefficients. For HRSTEM, the Cs-correction of the condenser lens leads to an improvement on the achievable spatial resolution due to the probe size reduction and the effective beam current increase. In addition to the instrumental TEM improvements, the development of more accurate models to describe the image formation and to support the TEM image simulation [15] provides a more reliable interpretation of experimental data. This work presents an overview on the HRTEM techniques state of art with several examples which indicate the development of such characterization tools towards the quantitative high resolution analysis of nanocrystalline materials.

2 HRTEM and HRTEM image simulation HRTEM image formation is based on the incident electron beam scattering by a thin sample. The technique is especially relevant for very thin crystalline samples, where the interaction of the several diffracted beams forms an interference patterns. As the diffracted components correspond to particular oriented periodic spacings on the sample, their interference may appear as a 2D periodic image which cannot be directly interpreted. The direct interpretation unfeasibility is mainly related to the HRTEM interference pattern dependency with the sample thickness and with the microscope configuration, especially the defocus value. However, the use of HRTEM image simulation can provide information about the sample crystalline arrangement. HRTEM image simulation is a crucial step in HRTEM characterization, which supports the atomic structure analysis. Multislice method [16] is the most widely used image simulation procedure and consists in the electrons wave function calculation after the electron beam interaction with the sample projected potential. Imaging distortions due to the lens aberrations, microscope instabilities and detection system defects [17] may be included in order to better reproduce the experimental imaging condition. The HRTEM imaging and image simulation combination has provided numerous examples of successful materials characterization. Remarkable examples can be noticed on the determination and refinement of crystalline WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


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structures [18] and the quantitative evaluation of the chemical composition profile along the thin films interface [19]. Figure 2 illustrates the combination of HRTEM imaging and image simulation on the evaluation of an anomalous anisotropic growth mechanism for SnO2 nanocrystals [20]. The used approach provided an unambiguous characterization of the growth mechanism and its preferential direction, which would not be feasible from direct imaging only.

Figure 2:

(a) SnO2 nanocrystal HRTEM image with an indexed FFT inset and a HRTEM multislice simulation (red square). (b) oriented attachment along the [110] is identified as the main growth mechanism. [From ref. [20], copyright © 2011 by RSC Publishing, reprinted with permission of authors.]

In addition, the accurate determination of nanocrystals 3D morphology by HRTEM simulation of nanocrystals [21] can be used to indirectly extract quantitative dopant segregation information [22] from systems where conventional analytical techniques are impracticable due to several experimental restrictions. Figure 3 depicts the 3D morphological modeling for Sb:SnO2 nanocrystals with different dopant contents. The nanocrystals models were compared to the Wulff [23] constructions based on ab initio surface energy calculations, providing the dopant atoms segregation for individual particles and its dependency with the doping level [24]. The presented examples depict the synergism between the HRTEM imaging and theoretical methods such as HRTEM image simulation and ab initio surface energy calculations. Therefore, this combination presents a remarkable potential as a tool for unveiling important nanocrystalline systems features, such as 3D morphology, dopant segregation, surface energy distribution, and growth mechanisms.



Figure 3:

(a), (e) Trimetric and (b), (f) [111] zone-axis oriented views for the Sb-doped nanocrystals geometric model. (c), (g) Original HRTEM images and (d), (h) superimpositions with the simulated HRTEM images. [Adapted from ref. [22], copyright © 2009 by ACS Publications, reprinted with permission of authors.]

3 Focal series reconstruction and geometric phase analysis Although the HRTEM image simulation can support accurate nanocrystals 3D morphology analyses, it is not always feasible to use such procedure to precisely evaluate the position of individual atoms or atomic columns in crystalline structures. To extract information which is directly related to the atomic positions, such as strain/stress state, a refinement on the HRTEM analysis is needed. Focal Series Reconstruction (FSR) [25] is a proposed calculation which is applied for the electron wavefunction restoration on the HRTEM imaging process and allows obtaining the sample projected potential without the interference of the microscope optical aberrations. The FSR implementation is based on the acquisition of a set of images under different objective lens defocus conditions, which is subsequently submitted to a restoration procedure that estimates the contrast transfer function (CTF) and the aberration coefficients implicated on the imaging process. A number of successful uses of FSR aiming a straightforward image interpretation [26], the correction of residual aberration [27], and the improvement of the HRTEM technique spatial resolution [28] have been already described in the literature. An example which illustrates the image improvement when FSR method is applied to structural analysis is depicted on Figure 4. These images show the comparison between an original HRTEM image and a phase image of the reconstructed wavefunction [29]. Geometric Phase Analysis (GPA) [30] is a methodology for measuring and mapping structural displacement fields on HRTEM images using a reference lattice. Its application to HRTEM image analysis allows the local lattice distortion evaluation, which can be directly related to the local strain state. This WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


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method has been effectively employed to study strain fields in semiconductors and metals [31]. The GPA implementation is based on the reciprocal space evaluation of a HRTEM image with respect to a reference undistorted lattice image, either obtained from a different region in the analyzed sample or from an image simulation procedure. As a spatial frequency spread around the characteristic lattice parameters frequencies is existent on strained samples, the determination of lattice distortions can be performed by the comparison with a reference image. The distorted components can be extracted on the reciprocal space and further translated to real space components through an inverse Fourier transform. Such analysis can be performed for two non-collinear spatial frequencies, resulting in a projected 2D distortion description of the sample which can be related to its 2D strain/stress state.

Figure 4:

(a) Example of an original HRTEM image and (b) the output from a FSR restoration from a CoSi2 sample. The insets exhibit a higher magnification detail with the crystal structure superimposition. [Adapted from ref. [29], reprinted with permission of authors.]

The GPA application is reported on the literature as an efficient tool to calculate strain maps on microelectronic components [32] and for the strain state analysis on individual nanocrystals [33]. A typical output from a GPA analysis is the distortion map from the analyzed HRTEM image, as exemplified on Figure 5, which includes the perpendicular, parallel and rotational strain components. An example of the combined user of FSR and GPA is presented on Figure 6 [34]. Cross-section HRTEM images of epitaxially grown Si1-xGex:Si(001) alloyed islands were used to characterize in a self-consistent way the strain configuration and the local chemical composition with high spatial resolution. By the combination of projected 2D chemical composition mapping from two different cross-sectioned zone-axis views, [100] and [110], it was possible to infer the chemical composition in a three-dimensional fashion [35].



Figure 5:

(a) Original Cs-corrected HRTEM image and (b) distortion mapping from a single dislocation on a CeO2 sample.

Figure 6:

(a) FSR reconstruction from a Si-Ge island including a higher magnification detail on the inset. (b) A simplified 3D chemical reconstruction was obtained after the strain chemical mapping of several Si-Ge islands at different zone-axis. An example from a single island: (c) chemical composition map, (d) parallel and (e) perpendicular strain projections. [Adapted from refs. [34] and [35], copyright © 2009 by ACS Publications, reprinted with permission of authors.]

The successful application of this technique to quantify the elastic behavior of the Si-Ge:Si(001) system shows that this methodology arises as a remarkable tool for accurate chemical and elastic state evaluation, which can be applied to several strained alloyed nanostructures, such as epitaxial islands, nanowires, nanocrystals, and thin films.



49

4 A novel approach for HRTEM imaging The improvement of HRTEM characterization depends on the comprehensive understanding of the underlying image formation theory. An important imaging condition representation is given by the Contrast Transfer Function (CTF), which represents how the spatial frequencies are transmitted through microscope imaging system. Since the 40’s, microscopists have been trying to improve the CTF characteristics by tuning the instruments parameters during the experiments or by the posterior treatment of the images. The main goal in this sense is to extend information transfer for high spatial frequencies. The most remarkable improvements were achieved by using induced defocus values on the objective lens in order to compensate its inherent positive spherical aberration coefficient [36]. However, a ground-breaking evolution happened on HRTEM with the Cscorrection possibilities. At a first glance, the use of Cs-correction adds an additional degree of freedom to the microscope tuning, allowing the CTF direct improvement for high spatial frequencies and improving the microscope information limit. A comparison between the HRTEM images from the same sample obtained by conventional and Cs-corrected microscopes, which is presented on Figure 7, clearly shows the image enhancement provided by the aberration corrector hardware.

Figure 7:

Raw HRTEM images from the same CeO2 nanocrystal from a (a) standard TEM microscope (1.25 Å information limit) and from a (b) Cs-corrected microscope (0.8 Å information limit).

A detailed investigation [37] of the electron scattering on the dynamical regime pointed out that non-linear contributions would enable the achievement of enhanced contrast transfer for low-Z elements by the use of negative Cs values. The Negative Cs Imaging (NCSI) [14] aims the contrast maximization at high frequencies by the microscope parameters optimization, including defocus and Cs adjustments. The NCSI application allowed some unique measurements WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)

50 Materials Characterisation V including the high resolution assessment of light elements atomic columns information, such as occupancy next to crystalline defects [38] and sub-angstrom displacement in magnetic domain walls [39]. Examples of NCSI HRTEM characterization for CeO2 sintered and nanocrystalline samples are depicted on Figure 8. Although the contrast for Oxygen columns is observed in both images, the direct analysis of a bulk-like sample HRTEM image (Figure 8a) is unfeasible due to a combination of factors. As the accurate thickness estimation is not achievable by HRTEM image simulation for this specific combination of crystalline system and imaging conditions, the intensity distribution cannot be directly associated to the Oxygen columns scattering. A solution for this issue is the sample projected potential assessment by the FSR application on NCSI HRTEM images (Figure 8b).

Figure 8:

(a) NCSI HRTEM image from a CeO2 bulk-like sample on the [100] ZA and (b) FSR from NCSI HRTEM images from a CeO2 nanocrystal on the same ZA orientation including a unit cell superimposition.

5 Concluding remarks State of art HRTEM techniques indicate that their use in conjunction with theoretical procedures for HRTEM image simulation and data analysis represents a vital tool for the quantitative evaluation of nanostructured systems. The depicted examples illustrate how this technique can simplify, or even present completely novel approaches, a number of materials characterization challenges where a quantitative high resolution assessment is required. The current development of HRTEM instrumentation and related methodologies indicates that its importance on materials characterization will grow on this decade, especially due to the novel possibilities achieved by aberration correction implementation and by on-growing combined use with HRSTEM imaging and spectroscopy techniques.



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Acknowledgements The authors acknowledge the financial support of the Brazilian research funding agencies FAPESP and FINEP; the German research funding agency DAAD. The authors would also like to thank the Ernst Ruska Centre staff at the Forschungszentrum Jülich and the QFA staff at the Universitat Jaume I for the fruitful discussions and scientific support.

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Ca(OH)2 nanoparticle characterization: microscopic investigation of their application on natural stones V. Daniele & G. Taglieri Department of Chemistry, Chemical Engineering and Materials, University of L’Aquila, Italy

Abstract Owing to conversion of lime into calcium carbonate, lime is usually adopted for conservative surface treatments. However, some critical aspects concerning the treatments reduced penetration depth, the binder concentration and the incomplete lime carbonation process still represent undesired limits and hindrances. In order to improve lime treatments, Ca(OH)2 particles with nanometric dimensions (nanolime) have recently been introduced in Cultural Heritage conservation (frescoes, stuccoes, ..). The aim of the present work is to characterize Ca(OH)2 nanoparticles synthesized by a chemical precipitation process starting from two base supersaturated aqueous solutions of calcium chloride and sodium hydroxide. After several washes, necessary to remove the sodium chloride, the aqueous medium is partially substituted by 2-propanol to improve the suspension stability; an alcoholic nanolime suspension is obtained, characterised by a given concentration and a residual water content. In order to identify the structure of the formed phases and the particles reactivity, the obtained nanolime is characterised by X-ray diffraction (XRD) and profile analysis; scanning and transmission electron microscopy (SEM-TEM) are performed too. The results show hexagonally plated and regularly shaped particles with side dimensions equal to or less than 300nm; moreover particles have pure crystalline features and a high reactivity in terms of the carbonation process. Finally, the Ca(OH)2 nanoparticles are applied on some natural lithotypes; SEM analyses are performed to evaluate penetration depth and grain adhesion of the nanolime treatments itself. From SEM micrographs a partial filling of the lithotypes pores, located at a distance from the surface of more than 200µm, is observed. Keywords: calcium hydroxide, consolidation, lime, nanoparticles, protection. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line) doi:10.2495/MC110051


1 Introduction Thanks to conversion of lime into calcium carbonate, lime water and milk are usually adopted for conservative surfaces treatments. In particular, lime water consolidation is generally obtained spraying the lime solution on the cleaned surface. To reach a good penetration, the treatment is repeated several times until the surface is able to absorb lime water [1]; some Authors indicate that could be necessary to repeat the application for 30-40 times [2]. Lime milk is used on the same basis as lime water [3]; nevertheless it involves greater amounts of lime with the same water volume, reducing the water percentage brought to the stone. However, some critical aspects concerning the treatments reduced penetration depth, the binder concentration and the incomplete lime carbonation process still represent undesired limits and hindrances. In order to improve lime treatments, Ca(OH)2 particles with submicrometric dimensions (nanolime) are recently introduced in Cultural Heritage conservation. Lime nanoparticles present the following advantages in stone, mortar and plaster consolidation: the possibility to penetrate deep into damaged zones (no limitations due to the particle size), high reactivity and fast reactions (such as carbonation) in the treated zones, high purity and defined composition [4]. Nanolimes are successfully employed on mural paintings, stuccoes and frescoes [5–9]; refurbishments of architectonical surfaces are considered too [10–12]. Lime nanoparticles are typically produced by a chemical precipitation process in supersaturated aqueous solutions of the reactants (calcium chloride and sodium hydroxide). To improve nanolime particles dispersion, the use of alcoholic solutions in place of aqueous ones is adopted; in fact, when 2-propanol alcohol is used as a solvent, dispersions of calcium hydroxide particles show a slower rate of agglomeration (and therefore, slower sedimentation rates) in comparison to aqueous media. This reduces the tendency for a white film to form on surfaces to be consolidated [13]. In this paper a nanolime suspension, characterised by a residual water content in the precipitated phase, is synthesised. X-ray diffraction (XRD) measurements and profile analysis are performed to characterise the nanolime samples and to evaluate the carbonation efficiency (yield). To correlate the produced nanolime to its properties, morphological characterisation is performed by scanning and transmission electron microscopy (SEM and TEM). Finally, the nanolime suspension is applied by brush on cleaned stone surfaces of natural lithotypes; in particular, the suspension is used in a diluted form (concentration of about 1mg/ml) in order to avoid the risk of leaving a white film on treated materials surfaces [11]. Several lithotypes, largely employed in architectonic fields, are considered: “Travertine”, “Pietra di Lecce”, “Pietra Serena” and “Basalto”. A local calcareous lithotype, “Poggio Picenze” is investigated too. SEM analyses allow one to investigate how the nanolime penetrates inside the stone porosity, in terms of the reached depth and the grains adhesion of nanolime treatments itself.



57

2 Experimental section Calcium chloride dihydrate (CaCl2·2H2O), sodium hydroxide (NaOH) and 2propanol pro analysi products, supplied by Merck, are used without further purification. Water is purified by a Millipore Organex system (R ≥ 18 M cm). Two aqueous solutions of 400ml, containing 0,3mol/l of CaCl2 and 0,6mol/l of NaOH respectively, are prepared. The NaOH solution (used as precipitator) is added dropwise into the CaCl2 solution (speed ≈ 4ml/min, temperature of 90°C). After about 24 hours two distinct phases are observed: a limpid supernatant solution and a white precipitated phase (NLW sample). In order to remove the NaCl produced, several deionised water washings are performed. Subsequently, in order to improve the suspension stability, the water content is partially substituted by 2-propanol; the obtained suspension is characterised by a water/2propanol ratio (W/A) of 0,75 and a final concentration of about 10mg/ml (NLA sample). SEM analyses (Philips XL30CP) are performed, depositing 0,2ml of the suspension on the specimen; TEM investigations (Philips CM100) are carried out dispersing 0,2 ml of the suspension in 50 ml of 2-propanol and depositing the sample on the suitable grid. As concerns XRD measurements, the sample is prepared maintaining the nanolime preparation for 20’ in ultrasonic bath (US) and then depositing 0,2ml of the suspension on a silica sample holder; measures are performed on dry sample, in laboratory conditions (T=20°C, relative humidity RH=40%). Each experimental diffraction spectrum is elaborated by a Profile Fit Software (Philips PROFIT v.1.0) and each crystalline phase is attributed by JCPDS patterns; the ratio between the CaCO3 peaks area and the spectrum total area is assumed as the carbonation process efficiency (yield). 2.1 Results and discussions TEM micrographs, obtained on NLA sample, are reported in Fig.1; in particular, hexagonally and regularly shaped particles of side dimension ranging from 100 to about 300nm are observed. Moreover, in Fig.1b) it is possible to note that all the particles appear overlapped in an ordered way and are so thin to be transparent to the electron beam.

a)

b)

Figure 1:

TEM micrographs on NLA sample: a) scale bar 500nm; b) scale bar 200nm.


58 Materials Characterisation V In Fig. 2, SEM images on NLA sample are shown. The particles, characterised by a prismatic features are recognizable to calcium carbonate of side dimension less than 500nm (Fig.2a). In Fig.2b) a perfectly hexagonal and regular shaped particle (marked as A) is observed, characterized by a side dimension less than 250nm.

A

a)

b) SEM micrographs on NLA sample (scale bar 1µm).

Figure 2:

XRD pattern on NLA sample shows the presence of Ca(OH)2 and CaCO3 phases (84-1276 and 85-1108 JCPDS patterns, respectively) with a corresponding yield value of about 80% (Fig.3a). On the contrary, XRD results obtained on NLW sample show as the corresponding yield decreases, reaching a value of about 50% (Fig.3b). The NLA higher yield values can be attributed to the presence of 2-propanol, that tends to “disagglomerate” the nanolime particles; so, a greater specific surface of the Ca(OH)2 particles exposed to air, can leads to a better carbonation process. *

* CaCO3 ° Ca(OH)2

* CaCO3 ° Ca(OH)2

° *

°

°

° *

a)

°

*

*

*

°

°

°

*

*

*

*

b) Figure 3:

XRD patterns: a) NLA sample; b) NLW sample.

3 Treatment on natural lithotypes: characterisation and results The considered samples, largely employed in the Italian historical architecture, are natural calcareous stones - Travertine, Pietra di Lecce and Poggio Picenze (the last one is a local lithotypes), a national sandstone - Pietra Serena - and an igneous lava stone of volcanic origin - Basalto. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


59

The nanolime treatment is carried out by brush, in laboratory conditions (T=20°C; R.H. 50%), applying the diluted alcoholic suspension (characterised by a concentration of 1mg/ml) on the dry and clean stones surface. In particular, 1ml of the suspension is applied on a side (5*5cm2) of the stone surface; when the solvent is evaporated, this procedure is repeated for 100 times (giving about 100mg Ca(OH)2 for each stone). For SEM investigations, each stone sample is broken along a plane perpendicular to the treated side and the section surface is observed. On the following figures, SEM micrographs, referred to untreated and treated samples section, are shown; in particular, in order to evaluate the treatment penetration inside the stone, pores situated at different surface distances are considered. 3.1 Travertine In Fig. 4 SEM micrographs on untreated Travertine stone sample is reported; in particular, a typical region, containing macropores and areas with a compact matrix is shown. For what concerns the treated stone, it is possible to note as the nanolime treatment covers the internal stone pore without filling it completely (Fig.5). In particular, the treatment, well recognisable in SEM micrograph referred to a pore at 150μm from the surface (Fig.5a-b), is found also in pores located at about 1mm from the surface itself (Fig.5c-d). This result can be attributed to the Travertine porosimetric structure, constituted by macropores, by which the nanolime treatment can be carried inside the stone.

A

a)

b)

Figure 4:

SEM micrographs on untreated Travertine stone: a) sample section (scale bar 500μm); b) zoom view on A pore (scale bar 100μm).

3.2 Lecce stone Pietra di Lecce is an organogenic marl calcareous stone, compact, generally porous and characterised by a fine-grained structure [14]. In Fig. 6, SEM micrographs on untreated sample are reported; it is possible to note the granular stone structure, constituted by various size clasts and by an intergranular porosity (Fig.6b).


60 Materials Characterisation V B

B

C a)

b) C

C

c)

d)

Figure 5:

SEM micrographs on treated Travertine stone: a) sample section (scale bar 200μm); b) zoom view on B pore localised at 150μm from the surface (scale bar 10μm); c)-d) different zoom views on C pore at about 1mm from the surface (scale bar 10μm and 5μm, respectively).

A

a)

b)

Figure 6:

SEM micrographs on untreated Lecce stone: a) sample section (scale bar 500μm); b) zoom view on A pore (scale bar 10μm).

Considering the treated sample, the treatment covers homogeneously the internal stone pores without filling them completely (Fig.7b-c). In particular, the treatment is also well recognisable in the pore located at 800μm from the surface, underlining the penetration depth reached from the treatment itself (Fig.7c).



61

B

B

C B a)

b) C

c) Figure 7:

SEM micrographs on treated Lecce stone: a) sample section (scale bar 200μm); b) zoom view on B pore localised at 250μm from the surface (scale bar 10μm); c) zoom view on C pore at about 800µm from the surface (scale bar 10μm).

3.3 Poggio Picenze stone Poggio Picenze lithotype, a local calcareous stone, is constituted by heterogeneous grained material; the stone surface and its internal structure are characterized by small-medium size pores and vacuoles (Fig.8) [14].

A

a)

b)

Figure 8:

SEM micrographs on untreated Poggio Picenze stone: a) sample section (scale bar 500μm); b) zoom view on A pore (scale bar 20μm).


62 Materials Characterisation V SEM micrographs referred to treated stone, show as the nanolime treatment tends to fill the internal cavity of the material, adhering to its grain, without occlude the pores. In particular, the treatment, recognisable in a pore located at a distance of 300μm from the surface (Fig.9), is not found in more internal cavities. B

B a)

b) B

c) Figure 9:

SEM micrographs on treated Poggio Picenze stone: a) sample section (scale bar 200μm); b)-c) different zoom views on B pore at about 300µm from the surface (scale bar 10 and 5μm, respectively).

In particular, in Fig.9b-c) internal stone cavities, that can be related to intergranular space, don’t appear covered by the precipitated particles; this fact can indicate that the nanolime treatment doesn’t modify the porosimetric system of the lithotype. 3.4 Pietra Serena stone This lithotype is different from those discussed above; in fact, it is a sandstone characterised by clay fractions. Moreover, this stone is constituted by different grains size, ranging from medium-fine to coarse ones (Fig.10) [14]. SEM micrographs referred to the treated sample are shown in Fig.11; the internal voids appear filled, on the walls, by precipitated particles (shaped like a small “flakes”) that tend to adhere to the clasts edges, without occluding the pores system (Fig.11b-c).



63

A

a)

b)

Figure 10:

SEM micrographs on untreated Pietra Serena stone: a) sample section (scale bar 500μm); b) zoom view on A pore (scale bar 20μm). B C

B a)

b) C

c) Figure 11:

SEM micrographs on treated Pietra Serena stone: a) sample section (scale bar 200μm); b) zoom view on B pore localised at 750μm from the surface (scale bar 10μm); c) zoom view on C pore at about 450µm from the surface (scale bar 10μm).

3.5 Basalto stone Basalt lithotype is an igneous rock lava of volcanic origin, typically composed by gray-black silicates [14]. It is an hard and resistant stone characterised by many pores and voids, of varying sizes, usually medium-large ones (Fig.12). WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


A

a)

b)

Figure 12:

SEM micrographs on untreated Basalto stone: a) sample section (scale bar 500μm); b) zoom view on A pore (scale bar 20μm). B

B

C a)

b) B

C

c)

d)

Figure 13:

SEM micrographs on treated Basalto stone: a) sample section (scale bar 200μm); b)-c) different zoom views on B pore localised at 500μm from the surface (scale bar 10 and 5μm, respectively); d) zoom view on C pore at about 700µm from the surface (scale bar 5μm).

SEM micrographs on the treated stone, show as the nanolime treatment tends to fill the internal cavity of the material, adhering to its grain, without occluding the porosimetric system. In particular, the treatment, is not found in pores more internal than 700μm.



65

4 Conclusions Nowadays lime nanoparticles are synthesised and employed for protection and superficial consolidation of several artworks such as stones, plasters, frescoes, wall paintings and paper documents. The obtained Ca(OH)2 nanoparticles are hexagonally plated and regularly shaped, with side dimensions generally less than 300nm, as shown by TEM observations. From XRD analyses, the carbonation process efficiency of the aqueous nanolime is about 50%, while the use of 2-propanol improves it reaching a yield value of about 80%. This result underlines the 2-propanol role in disagglomerating the particles, leaving a higher specific surface exposed to air. SEM analyses are performed to evaluate the interaction between the nanolime treatments and the porosimetric system of some natural lithotypes. The images show the partial filling of the pores in all the examined lithotypes, without occluding them. On the contrary, considering all the lithotypes, the treatment penetration is observed at a limit distance from the surface, ranging from 300µm (Poggio Picenze stone) to about 1mm (Travertine stone).

References [1] Vinardi, M.G., Cully, M.H., Brunetto, A., La reversibilità nel restauro. Riflessioni, esperienze, percorsi di ricerca, In: XI National Congress Scienza e Beni Culturali: “La Pulitura delle Superfici dell’Architettura”, Bressanone, XIX, pp. 399-405, 2003. [2] Sten, P., Lime Water Consolidation, ICCROM, 53-61, 1981. [3] Turriziani, R., I leganti ed il calcestruzzo, Edizioni Sistema, Roma, pp. 70-82, 1972. [4] Pianski, J., Brümmer, K., Ziegenbalg, G., Nano-particles for stone conservation–state of the art, characteristics and recent developments, In: Stonecore – “Recent progress in the consolidation of calcareous materials”, Litomyšl, Czech Republic, Aprile 2010. [5] Baglioni, P., Cesari, C., Dei, L., Giorgi, R., Grassi, R., Lorenzetti, M., Mauro, M., Pinzani, D., Ruffo, P., Schonhaut, G., Stucco forte veneziano con “Tecniche della Tradizione”: proprietà chimico-fisiche, degrado in ambiente lagunare e conservazione mediante nanofasi cristalline di Ca(OH)2, In: National Congress Scienza e Beni Culturali: “Lo Stucco. Cultura, Tecnologia, Conoscenza”, Bressanone, XVII, pp. 289-298, 2001. [6] Ambrosi, M., Baglioni, P., David, P.R., Dei, L., Giorgi, R., Lalli, C., Lanterna, G., Mairani, A., Matteini, M., Rizzi, M., Schonhaut, G., Inorganic consolidants and protectives for architectonic surfaces: experimental tests on Santa Prisca Church Apse in Rome, In: 2nd International Congress on “Science and Technology for the safeguard of Cultural Heritage in the Mediterranean Basin”, Vol. 2, pp. 873-877, 1992. [7] Dei, L., Salvadori, B., Arlango, E., Pietropoli, F., Scardellato, C., Gli affreschi del XIII e XIV secolo nella Cripta di San Zeno a Verona: la sperimentazione della nanocalce dispersa in alcol iso-propilico durante WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


[8]

[9]

[10]

[11]

[12]

[13]

[14]

l’intervento conservativo, In: International Congress Scienza e Beni Culturali: “Sulle pitture murali. Riflessioni, conoscenze, interventi”, Bressanone, XXI, pp. 293-302, 2005. Dei, L., Radicati, B., Salvadori, B., Sperimentazione di un consolidante a base di idrossido di calcio nanofasico sugli affreschi della Cappella del Podestà al Museo del Bargello di Firenze: aspetti chimico-fisici e prove di colore, In: International Congress Scienza e Beni Culturali: “Sulle pitture murali. Riflessioni, conoscenze, interventi”, Bressanone, XXI, pp. 99-108, 2005. Dei, L., Favaro, M., Giorgi, R., Portieri, R., Gli stucchi della Cappella della Madonna di Lourdes nella Chiesa di S. Giovanni Evangelista in Venezia: caratterizzazione del degrado e pre-consolidamento mediante dispersioni alcoliche di calce, In: National Congress Scienza e Beni Culturali: “Lo Stucco. Cultura, Tecnologia, Conoscenza”, Bressanone, XVII, pp. 557-566, 2001. Slížková, Z., Drdácký, M., Moreau, C., Frankeová, D., Nosál, L., Consolidation of porous limestone with suspensions of calcium hydroxide nano-particles in alcohols, In: Stonecore – “Recent progress in the consolidation of calcareous materials”, Litomyšl, Czech Republic, 2010. Daniele, V., Taglieri, G., Nanolime suspensions applied on natural lithotypes: the influence of concentration and residual water content on carbonatation process and on treatment effectiveness, Journal of cultural heritage, 11, pp. 102-106, 2010. Ambrosi, M., Giorgi, R., Salvadori, B., Dei, L., Nanotecnologie per la conservazione di affreschi e materiali lapidei a base carbonatica, In: 1st International Workshop on Science, Technology and Cultural Heritage, Venezia, pp. 29-33, 2004. Hansen, E., Doehne, E., Fidler, J., Larson, J., Martin, B., Matteini, M., et al., A review of selected inorganic consolidants and protective treatments for porous calcareous materials, Rev Conserv., 4, pp. 13–25, 2003. http://www.italithos.uniroma3.it/index.php.



67

Nanocarbon composite materials with optical response on radioactive waste M. Vantsyan1, G. Popova1, E. Karpuzova1, M. Bobrov1, O. Plaksin2 & E. Dabek3 1

D. Mendeleyev University of Chemical Technology of Russia, Russia A.I. Leypunsky Institute for Physics and Power Engineering, Russia 3 Science & Technology Branch Environment Canada, Ottawa, Canada 2

Abstract Nanocarbon materials have numerous unique features – high porosity, large specific surface area, chemical inertness, radiation stability, etc. We applied nanocarbon/nanodiamond and silicon carbidecomposites as matrix for optical chemochips construction. Composite elements consist of porous nanocarbon substrate with specific chromophores introduced to nanosize pores and siliconorganic coating. Similar multicomponent composition has a response to radiation, in particular, -irradiation. By using diarylethenes as sensitive chromophores, their electronic and/or luminescent spectra data may be applied for doze power detection. Quantum chemistry methods, computer simulation were considered for optimal design of nanocarbon/organic chromophore hybrid. Experimental data and modeling have shown that diarylethenes are able to change color under -irradiation in molecular crystal phase only. Weak interactions (inter-, intramolecular and binding with hydrogen-containing walls in pores) play key role under irradiation. Induced self-organization in limited volume is considered. SWAXS and AFM data are discussed. Nanodiamond composition possesses luminescent response to -, - and -irradiation. Siliconorganic polymer and SiC composition are neutral to irradiation, they are stable in extreme conditions. Application of composite elements with optical sensing as multifunctional chemochip fragment for atmospheric media monitoring is discussed. Keywords: nanocarbon, nanodiamond, silicon, composite, chemochip, chromophore, radiation, computer simulation, optical response, nuclear waste.



1 Introduction One of the most important directions of modern nanoscience and nanotechnology is the so-called sensor technology combining the latest achievements in highly sensitive intelligent materials, first of all, multifunctional chemochips with immobilized sensing fragments with nanoscale function [1–4] and different microdevice construction with high processing speed and reliability. Of especial interest are the materials with high sensitivity and adaptability [5]. In this context, great attention is being given to the selfassembling systems capable of hierarchical ordering from the nano- to mesoscopic level [6], which may lead to elaboration of intelligent, in particular, optically sensing, materials responsive to weak external factors such as temperature, irradiation, pH, pollutants in the environment, etc. Typically, these devices (chemochips) consist of solid inorganic support and sensing fragments immobilized on it. Nanocarbon materials are promising inorganic supports as they have a large variety of unique and specific properties – high porosity, large specific surface area, chemical inertness, radiation stability and others [7]. To date, highly porous (up to 40% vol.) composites can be prepared on the basis of nano-sized diamond. Diarylethenes seem to be the most suitable objects to be used as sensing units. These are photochromes changing from colourless to coloured under irradiation; normally, the colour disappears upon UV or visible-light irradiation [8]. Diarylethenes have numerous advantages such as thermal stability, linear dose dependence of the coloration, compared to other photochromic compounds. On the other hand, diarylethenes exhibit their sensing properties in highly ordered state (e.g. in single crystal). In this context, of great importance is prediction of diarylethene molecules self-organization pattern. Computer modeling and quantum chemistry calculations help greatly to successfully solve applied science problems of predicting optimal synthetic pathways, structure and properties of materials resulting in decrease in time and costs of initial laboratory investigations. For integrated solution of the problem it was necessary to predict diarylethenes self-organisation in a limited volume i.e. to estimate the possibility for chromophores introduction into porous composite and to evaluate the assembly parameters. In this study, we applied experimental techniques together with a set of up-todate computing chemistry methods, Cambridge Structural Database and specialized software. We obtained models of diarylethene molecules and their assemblies; non-covalent interactions providing formation of 2D and 3D supramolecular clusters and crystal structures were revealed.

2 Results and discussion In this study, for multifunctional detector device creation we used efficient adsorption properties of nanoporous carbon composites nanocarbon/nanodiamond and silicon carbide [9, 10] subsequently covering the composite matrix with polymeric photochromic composition. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


69

As nanodiamond composite matrix is black it is reasonable to introduce only luminescent molecules whereas silicon carbide composites are suitable for colour-changing chromophores. Therefore, we applied diarylethenes that change both luminescence and colour under different types of irradiation. Moreover, diarylethenes (DAE) have a unique property to change their colour under radioactive irradiation and to reversibly restore it under UV-irradiation. We have chosen a diarylethene – 1,2-bis(2,5-dimethyl-3-thienyl)perfluorocyclopentene changing from colourless to coloured under -irradiation; the colour disappears upon UV or visible-light irradiation [11] (Fig.1). F

F

F F F

F F

S

H 3C

S

CH3 H3C

Open form colourless

Figure 1:

F F

F F

UV or visible-light irradiation

CH3 H3C

-irradiation

F

CH3 S H3C

S

CH3

Closed form coloured

Coloration under -irradiation and bleaching under UV or visiblelight irradiation.

Nanodiamond/nanocarbon composites samples were black tablets with diameter of 8 mm and thickness of 3 mm. They contained 28% vol. nanodiamond, 15% vol. graphite-like carbon with porosity of the material 57% vol. and pore size 8 – 10 nm (as found by capillary condensation method). Silicon carbide nanocomposite samples were grayish tablets with a diameter of 8 mm and thickness of 3 mm. Composites surface was studied by atomic force microscopy method (AFM); the average pore size was found to be up to 10 nm [10]. However, pore shape and depth was not estimated with high accuracy. Theoretically, different pores in composites can have conic and cylindrical shape. For calculations we used an approximation that all pores in composites are cylindrical with pore diameter not more than 10 nm. Average grain size of the composites was estimated by wide-angle x-ray scattering method (WAXS). For nanodiamond/nanocarbon composite it was 4 nm, for silicon carbide – 12.5 nm. Chromophores were introduced by impregnation method: composites were treated by 2% mass solution of diarylethene in chloroform with subsequent drying under vacuum. Luminescence spectra of modified nanoporous matrices exhibit a new emission band corresponding to starting diarylethene. The intensity of this new


70 Materials Characterisation V band is approximately 15–20% from intensity of the luminescence peak of starting diarylethene. A preliminary study has shown rather high radiation stability of nanocarbon composites , silicon-organic coating and chromophores. Diarylethene molecules are conjugated aromatic systems capable not only of hydrogen bonding but also of stacking interactions [12]. Pore size and shape also influence self-organization of chromophores when filling the limited volume. A molecular geometry modeling for diarylethene molecule and for its molecular aggregates was performed by semi-empirical quantum chemistry methods АМ1 by programme complex GAMESS [13]. The initial approximation of the molecule (derived from x-ray diffraction data) was taken from Cambridge Structural Database [14]. After optimizing the geometry parameters of diarylethene molecule, we have estimated maximal distances between atoms centres along the axes of three main inertias. These distances were ~9,9 Å (along axis 1), ~11,5 Å (along axis 2), ~6,6 Å (along axis 3) The optimization of geometry parameters of two diarylethene molecules assembly was performed by non-empirical quantum chemistry method GAMESS DFT 6-31G(d,p) [13]. These methods are also suitable for calculation of bond critical points, in particular, bond critical points of non-covalent interactions [15]. The size of the optimized model of the molecular aggregate including two chromophore molecules (Fig. 2) have also been estimated by maximal distances between atom centres along three axes. Two diarylethene molecules are arranged in two parallel planes and are bonded by weak bonds. One can assume two probable aggregate structures – one is formed by four СF…H bonds and the other – by two СS…H bonds. Both aggregates dimensions are close to one another – these are 19* 11* 9 Å. The proposed four chromophore molecules aggregate structure (Fig. 3) has maximal distance between atom centres along one of the axes, equal to 29 Å. The distance along two other axes is 18 Å and 9 Å. As a result of the calculation performed, it was found that two diarylethene molecules are also arranged in two parallel planes. In order to reveal weak noncovalent interactions in two diarylethene molecules assembly we have applied R. Bader quantum topological analysis of atoms in molecules [15]. By programme complex AIM-2000 bond critical points were calculated and a molecular graph was drawn for two diarylethene molecules assembly [16]. We have calculated bond critical points between carbon atoms, distances and angles of weak noncovalent interactions, local characteristics in critical points (Table 1). This calculation has elucidated weak bonds between fluorine and hydrogen atoms, these bonds are pairwise symmetrical. Moreover, weak bonds were found inside the molecule, by bonds fluorine – hydrogen an intramolecular cycle is formed.



71

(a)

(b) Figure 2:

Assemblies of two diarylethene molecules; (a) – formed by four СF…H bonds, (b) – by two СS…H bonds. Non-covalent bonds are shown by dotted line.

Table 1:

Distances, angles, and local parameters in critical points (3,-1) corresponding to weak interactions.

CP number 1 2 3 4 5 6 7 8 9

R А 2.54 2.53 2.53 2.54 2.60 2.60 2.78 2.78 4.20

Angle, degrees 139.5 146.8 146.8 139.4 108.5 108.5 138.3 138.3 -

Contact CH38…F45 CH38…F47 CH77… F8 CH77… F6 CH61…F47 CH22… F8 CH22…H70 CH38…C23 C21…C60

CP number 55 48 52 35 36 24 43 47 33

b

2b

gb

vb

Econt

a.u.

a.u.

a.u.

a.u..

kcal/mole

0.00636 0.00645 0.00645 0.00635 0.00677 0.00677 0.00195 0.00195 0.00117

0.02822 0.02825 0.02825 0.02822 0.03319 0.03320 0.00619 0.00619 0.004150

0.005769 0.005794 0.005794 0.005768 0.006470 0.006471 0.001114 0.001114 0.0007669

0.004483 0.004525 0.004525 0.004482 0.004642 0.004642 0.0006811 0.0006811 0.0004963

1.41 1.42 1.42 1.41 1.46 1.46 0.21 0.21 0.16



(a)

(b) Figure 3:

Assemblies of four diarylethene molecules: (a) – formed by two СS…H bonds and eight СF…H, (b) – by four СS…H bonds and four СF…H bonds.

Energies of the above mentioned interactions were estimated by local density of the potential energy (νb) in a critical point. For this, an empirical formula was used: E (kcal/mole) = 313.754•vb (atomic units) [17]. So, the overall energy of interatomic non-covalent interactions was Econt = -9.15 kcal/mole.

3 Conclusions Experimental techniques (e. g. AFM, x-ray scattering and others) applied together with a set of up-to-date computing chemistry methods aided greatly for subsequent creation of a chemochip sensitive to -irradiation based on WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


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nanocarbon composites. Computer modeling has revealed the self-organization pattern for diarylethene molecules in limited volume. Diarylethene molecule dimensions (as well as those for its aggregate of two and four molecules calculated by semi-empirical quantum chemistry methods) are estimated. By density functional theory geometry parameters in assembly of two diarylethene molecules were optimized. The quantum-topological electron density analysis has revealed weak non-covalent interactions fluorine-hydrogen, carbon-hydrogen, carbon-carbon. In separate diarylethene molecules a cycle formed by non-covalent interactions fluorine-hydrogen is found. The overall energy of interatomic non-covalent interactions was Eконт = -9.15 kcal/mole. Quantum chemistry and quantum topological analysis methods are promising tools for revealing interatomic interactions when studying formation of molecular assemblies at the nanoscale in pore filling modeling.

Acknowledgements The research is being supported by Ministry of Education and Science or Russian Federation (Project 2.2.2.2.325) and ISTC (Project #3891).

References [1] Al-Azzawi A. (ed). Photonics: Principles and practices, CRC Press: Boca Raton, 2007. [2] Davies A.G. and Thompson J.M.T. (eds). Advances in nanoengineering: Electronics, materials and assembly, Imperial College press: London, 2007. [3] Fryxell G.E. and Cao G. Z. (eds). Environment application of nanomaterials – Synthesis, sorbents and sensors, Imperial College Press: London, 2007. [4] Lu G.Q. and Zhao X.Z. (eds), Nanoporous materials: Science and Engineering, Imperial College Press: London, 2004. [5] Ong K. G., Yang X., Mukherjee N., Wang H., Surender S., Grimes C.A. A wireless sensor network for long-term monitoring of aquatic environments: Design and implementation, Sensor Letters, 2 (1), pp. 48-57, 2004. [6] Grimes C.A, Dickey E.C., Pishko M. V., Encyclopedia of Sensors, ASP Press: N.-Y., 2004. [7] Seki Y., Impact of low activation materials on fusion reactor design, J. Nucl. Mater, 258-263, pp. 1791-1797, 1998. [8] Irie S., Irie M., Radiation-induced coloration of photochromic dithienylethene derivatives in polymer matrices, Bull. Chem Soc. Jpn., 73, pp. 2385-2388, 2000. [9] Lisichkin G.V. (ed.) Chemistry of Grafted Surface Compounds, Fizmatlit: Moscow, 2003. [10] Gordeev S.K., Nanocarbon Materials, Nanotechnics (Rus.), 2005, pp. 3-11. [11] Morimoto M., Kobatake S. Irie. M., Photochromism of diarylethenes in nanolayers of a single crystal, Photochem. Photobiol. Sci., pp. 1088-1094, 2003. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)

74 Materials Characterisation V [12] Yamada T., Kobatake S., Muto K., Irie M., X-ray Crystallographic Study on Single-Crystalline Photochromism of Bis(2,5-dimethyl-3-thienyl) perfluorocyclopentene, J. Am. Chem. Soc., 122, pp. 1589–1592, 2000. [13] http://classic.chem.msu.su/gran/gamess/index.html. [14] CSDB, http:// www.ccdc.cam.ac.uk [15] Bader R., Atoms in Molecules: A Quantum Theory, Oxford University Press, 1994. [16] http://www.aim2000.de [17] Espinosa E., Alkorta I., Rozas I., Elguero J., Molins, E., Topological Analysis of the Electron Density Distribution in Perturbed Systems. I. Effect of Charge on the Bond Properties of Hydrogen Fluoride, Chem. Phys. Letts., 336, pp. 457-464, 2001,.


Section 3 Corrosion problems



77

Evaluation of the fretting corrosion mechanisms on the head-cone interface of hip prostheses I. Caminha1, C. R. M. Roesler2, H. Keide1, C. Barbosa1, I. Abud1 & J. L. Nascimento1 1

Laboratory of Characterization of Mechanical and Microstructural Properties, National Institute of Technology (INT), Rio de Janeiro (RJ), Brazil 2 Laboratory of Biomechanical Engineering, University Hospital, Federal University of Santa Catarina (UFSC), Brazil

Abstract Fretting-corrosion is one of the main concerns in the application of hip prosthesis. This type of information is very important in the stage of orthopedic implants design, with the purpose of minimizing the amount of tissue exposed to corrosion products which are released during the permanence of the prosthesis in the patient. The residual corrosion products of stainless steels are associated to the occurrence of several adverse reactions in the human body. The knowledge about these corrosion products is extremely important in the phase of the project of hip prostheses, aiming at the minimization of the amount of exposure of the organic tissues to corrosion products released during the permanence of the prosthesis inside the patient. In the present work the mechanical stability and fretting corrosion resistance of modular hip prosthesis, which was fabricated with ASTM F 138 austenitic stainless steel, were evaluated according to the criteria of ASTM F 1875 standard, method I, which prescribes long term test, with the purpose of determining the amount of damage through the quantification of the corrosion products and debris which resulted from the fretting corrosion conditions. The mechanical tests were performed in a servohydraulic mechanical testing machine and the modular interfaces were exposed to an electrolytic 0.9% NaCl in distilled water solution and subjected to a minimum load of 230 N and a maximum load WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line) doi:10.2495/MC110071

78 Materials Characterisation V of 2.3 kN, frequency of 5 Hz for 10 million cycles, according to ASTM F 1875 standard, thus simulating real conditions of use. Load ratio (R = 10) was determined in ASTM F1440 standard. A significant amount of fretting-corrosion products were observed: 0.22 g resulted from five specimens after 107 million cycles, and the diffraction X-Rays tests showed the presence of the crystalline phases Fe2O3, Fe3O4 and Cr2O6. It can be concluded that the mass loss in the head-cone connection allowed the entrance of the physiological saline solution in the inner region of the head, increasing the predicted micro movement in the head–cone interface, resulting in an accelerated process of fretting-corrosion, and consequently the liberation of debris and corrosion products that could lead to adverse biological reactions. Keywords: fretting corrosion, hip prosthesis, modular components, mechanical properties.

1 Introduction The increasing life expectancy in the last decades and trauma risk due to the more frequent exposure of individuals to different types of accidents and current insecure life conditions has led to an increased search for orthopedic implants with high performance, capable of resisting to even more severe loads for longer times. For this reason, metallic materials, which were developed for applications in orthopedic implants, must present some specific properties, such as biocompatibility, strength and resistance to degradation (by wear or corrosion). In spite of the increasing use of titanium and its alloys for this application in the last decades, in Brazil ASTM F138 austenitic stainless steel [1, 2] is still very used in the fabrication of modular components for total hip arthroplasty due to its lower cost, together with good mechanical properties and satisfactory corrosion resistance, since the main customer is Health Ministry, which supplies the demands of public hospitals in Brazil. The use of modular components in the total hip arthroplasty presents several advantages, such as a great variety of stem geometry, which allows a better choice of modular component for the surgeon, according to the patient disease, mainly in the case of revision surgeries, thus minimizing the risk of inadequate procedures. However, the modular interfaces are subjected to micro movements, which can result in fretting and corrosion, leading to the release of debris, which can cause adverse reactions and accelerated wear in the articulation interface. “Fretting” can be defined as a wear phenomenon which occurs when two solids in contact are subjected to tangential oscillatory movement with small displacement amplitude. It can be described more accurately as a movement whose amplitude is smaller than the contact extent. The presence of a corrosive environment contributes to accelerate the wear process and this condition is defined as “fretting-corrosion”. The occurrence of corrosion results in degradation processes which reduce the structural integrity of the implant and release products whose reaction can be harmful in contact with organic tissue.



79

When the prostheses are fabricated with ASTM F138 stainless steel, the generation of a passive barrier in the surface (passive layer) is the main obstacle to corrosion process. This film, which is composed by metallic oxides, formed spontaneously on the metal surface, and whose thickness can be increased with surface treatments, avoids the migration of metallic ions from the metal to the solution and the migration of anions from the solution to the metal through the metal-solution interface (physiological environment). For being effective barriers these films must be compact and cover completely the metallic surface, they must have an atomic structure which limits the migration of metallic ions and/or electrons through the metallic oxide/solution interface, and they must be capable of remain on the surface of these alloys even when subjected to mechanical and abrasive stresses in the clinical use of prostheses. The most frequently mentioned mechanisms of failure in these components are aseptic loosening, periprosthetic osteolysis and metallosis and are directly related to the debris released by the prosthesis components, mainly the non cemented prostheses [3]. According to the literature [4], the residual products of corrosion in stainless steel based implants are associated to tissue necrosis, inflammation, allergenic reactions and even cancer. The aim of the present work is an evaluation of fretting corrosion in modular prostheses fabricated with ASTM F 138 stainless steel, in order to minimize the adverse reactions during the permanence of the prosthesis in the patient.

2 Material and methodology In this work five modular hip prostheses, composed by heads and stems, five stems and five heads, fabricated with ASTM F 138 from the same batch, were analyzed. A SZX16 model Olympus stereomicroscope was used for comparing the surfaces of the modular components before and after fretting corrosion test. In Fig.1 an as received representative sample of the batch can be observed.

Head Figure 1:

Cone

Stem

Head and stem of the hip prosthesis fabricated with ASTM F 138 stainless steel.


80 Materials Characterisation V The fretting corrosion tests on the cone-head interface were performed according to method I of ASTM F 1875 standard, which prescribes a 0.9% NaCl physiological saline solution, thus simulating in vivo conditions, since that when implanted in patients these components are in contact with body fluids. With the purpose of analyzing both fatigue resistance of hip stems and mechanical stability of modular components the experimental procedures described in ASTM F 1440 [5] and ISO 7206-4 [6] were also adopted. These tests were performed in two universal machines for mechanical testing (Instron model 8872) with a 25 kN load cell with the following characteristics, described in ASTM F 1875 standard [7]: -

Maximum load: 3.3 kN;

-

Frequency: 5 Hz;

-

Test finish after 10 million cycles;

-

Error of applied load lower than 1% of the maximum load;

-

Control function of cyclic load (sinusoidal);

-

Monitoring of vertical displacement of the prosthesis head and registering of cycles counting;

-

Computerized operation;

-

Equipment for recirculation and heating of solution test with temperature control within 37ºC±1ºC. Before starting the test each modular component composed by head and stem was prepared according to the following procedure: - Assembling the modular component and checking the fittings; - Placing the modular component in a device which allows alignment of angles referred in the standard, in concern to the load direction; - Drying the mounting; - Introducing the environment chamber together with the metallic cup with the mounted prosthesis; - Adding the saline solution (0.9% NaCl in distilled water) to the environment chamber; - Starting the pump for recirculation and heating the test solution; - Fitting the complete assembly in the test machine (Fig. 2). The machine was programmed for the application of the load in the center of the prosthesis head, with minimum 0.3 kN and maximum 3.3 kN loads, with sinusoidal 5 Hz frequency according to the conditions specified in ASTM F1875 standard. Maximum load requirements are specified in ISO 7206-4-2010 standard. A limiting condition was imposed for automatically stop ping the machine: a vertical displacement of 5 mm. If this value is not reached the test is finished after completing 10 million cycles of loading and unloading. The parameters of sample placement follow the requirements of ASTM F1440 standard. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


Figure 2:

81

Complete assembly before starting test in INSTRON 8872 machine.

X-ray diffraction analysis was carried out in the following conditions: anode material: Cu Kα, angle between 10 and 100º, step size 0.05º, scan step time: 1 s, receiving slit size: 0.03 mm, current: 40mA, voltage: 40 kV.

3 Results and discussion The images of the modular surfaces (head and cone), which were obtained by optical microscopy, did not present surface defects such as cracks other types of flaws, introduced by the fabrication process, which could influence the results of this test. In Figs. 3 and 4 the aspect of the surface of the modular components (head and cone) of a hip prosthesis, before and after testing, are presented. The other prostheses have similar aspects. A significant corrosive attack inside the head and in the cone can be observed after 1 million cycles with increasing rates along the test, releasing corrosion products which resulted in mass loss in the head-cone connection. The mass loss, which was observed in the test solution, allowed the introduction of saline solution inside the head, increasing the predicted micro movements, leading to an accelerated wear (friction), with tearing of debris, according to Fig. 6. No type cracks and mechanical failure could be observed in the five hip prostheses, which were analyzed, thus meaning that the mechanical stability of the modular component was preserved along the test.

a

Figure 3:

b

Detailed view inside the heads: (a) before and (b) after the test. Magnification: 10X.



b

a

Head-cone interface

Figure 4:

Detailed view of the cone region: (a) before (b) after. Magnification: 10X.

Figure 5:

Detailed view of the head-cone interface connection shown in Fig. 4-b. Magnification: 60X.

Figure 6:

Detailed view of tearing which was observed on the cone. Magnification: 60X.



83

In Figs. 7 and 8 the corrosion products and particulate residues (debris), resulting from the fretting corrosion test can be observed in the bottom of the environment chamber (indicated by an arrow in fig. 8). These ones are consequence of the damages caused by fretting and corrosion fretting in the head-cone interface of the primary hip prosthesis. These images represent the five hip prostheses subjected to this analysis.

Figure 7: Test assembly after 106 cycles.

Figure 8: Detail.

Until 1 million cycles neither corrosion products nor particulate debris could be observed, but after this period the degradation of the modular component was increasing and localized on the head-cone, as can be seen in detail in Fig. 5. After 10 million cycles the corrosion products and particulate debris were quantified, resulting in an average value of 0.22 g. In the literature there is only one citation of fretting-corrosion products quantification, but in their experiment the cycle number was varied, thus hindering any type of comparison with the results presented in this work. For better comprehension of fretting corrosion mechanisms in hip modular components the variation of load and number of cycles is necessary. The characterization of the corrosion products and particulate residues by X-ray analysis is also very important, in order to identify the phases and type of metallic particles which were released, with the aim of minimizing the incidence of aseptic loosening, periprosthetic osteolysis and metallosis in the surgeries of total hip arthroplasty. In the present work a complete characterization of the fretting corrosion products was not concluded, but it can be observed that they presented a predominantly ellipsoidal shape with size distribution approximately uniform, whose larger diameter was around 500 μm. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)

84 Materials Characterisation V Anyway, the X-ray diffraction analysis performed in the corrosion products showed the presence of the following crystalline phases: Fe2O3, Fe3O4 and incipient Cr3O8, compatible with the ASTM F138 stainless steel.

4 Conclusion The evaluation of fretting corrosion on the head-cone connection interfaces of five modular hip prosthesis fabricated with stainless steel led to the following results:  A significant grade of localized corrosion in the inner region of the head and in the head-cone interface was observed, originated from the fretting corrosion and fretting during 10 million cycles, resulting in particulate debris and corrosion products.  In order to achieve a better comprehension of the fretting corrosion mechanisms, more tests are necessary, including changes in the applied load and number of cycles.  At first glance, fretting corrosion does not affect the mechanical stability of the modular components, at least until 107 cycles.

References [1] Cavalcanti, E.H. de S., Souza, S.M.C. de, Ferreira, C. de A., Campos, M.M., Abud, I. de C. and Palmeira, L., (2002), “Avaliação da Resistência à Corrosão de Prótese Total de Quadril de Aço Inoxidável Austenítico Removida de Paciente”, 22° CONBRASCORR – Congresso Brasileiro de Corrosão, Salvador – Bahia. [2] Giordani, E.J., Ferreira, I. and Balancin, O., (2007), “Propriedades mecânicas e de corrosão de dois aços inoxidáveis austeníticos utilizados na fabricação de implantes ortopédicos”, REM: R. Esc. Minas, Ouro Preto, 60(1): 55-62. [3] Crestani, M.V., Boschim, L.C. and Schwartsmann, C.R., (2004), “Metalose Simulando Tumor Abdominal” , Revista Brasileira de Ortopedia, Nov/Dez. 2004. [4] Merritt, K.; Brown, S.A.; (1985), “Biological Effects of Corrosion Products from Metals”, Corrosion and Degradation of Implant Materials: Second Symposium, ASTM STP 859, A.C. Fraker and C.D. Griffin, Eds., American Society for Testing and Materials, Philadelphia, 1985, pp. 195-207. [5] ASTM F 1440 – 92 (Reapproved 2008), Standard Practice for Cyclic Fatigue Testing of Metallic Stemmed Hip Arthroplasty Femoral Components Without Torsion. [6] ISO 7206-4-2010 - Implants for surgery - Partial and total hip joint prostheses -- Part 4: Determination of endurance properties and performance of stemmed femoral components. [7] ASTM F 1875 – 98 (Reapproved 2004), Standard Practice for Fretting Corrosion Testing of Modular Implant Interfaces: Hip Femoral Head-Bore and Cone Taper Interface. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


85

Improving corrosion performance by surface patterning M. Bigdeli Karimi, V. Stoilov & D. O. Northwood Department of Mechanical, Automotive and Materials Engineering, University of Windsor, Canada

Abstract Based on the idea that hydrophobic (low or non-wettable) surfaces can decrease the contact area between a corrosive solution and a surface, thereby potentially rendering the material more corrosion resistant, the effect of surface patterning on the corrosion behaviour of nickel was investigated. The surface patterning consisted of an array of holes of various diameters (D) and inter-hole spacings (L) that were produced by a laser ablation process. The corrosion behaviour of the patterned surfaces was studied using a potentiodynamic polarization method in a 0.5M H2SO4 electrolyte and compared with that of a polished reference sample. Following the potentiodynamic polarization corrosion test cycle, the corroded surfaces were examined using scanning electron microscopy (SEM) for morphological features and white light interferometry (WLI) to determine the surface roughness. The changes in surface morphology were related to the corrosion behaviour. A relationship was found between D, L, and the corrosion current density (Icorr), whereby the higher the (D/L )2 ratio, the higher the Icorr value. The corrosion potential (Ecorr) of all surface patterned samples was lower (less noble) than that of the reference sample in all tests. Keywords: corrosion resistance, surface patterning, hydrophobicity, laser ablation.

1 Introduction Wettability of solid surfaces is an important property and depends both on the surface chemistry and on the surface topology. Recently, hydrophobic and superhydrophobic surfaces (water contact angles larger than 150°) have received a lot of attention, due to their important applications ranging from self-cleaning WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line) doi:10.2495/MC110081

86 Materials Characterisation V materials and in-flight icephobic treatment [1] to microfluidic devices. The archetype superhydrophobic surface is that of the lotus leaf, on which a water droplet apparently forms a sphere, reducing the area of contact. This behavior, known as the lotus-leaf-effect, is found to be a result of the hierarchically patterned structure. Effectively the fluid cannot penetrate the air gaps between the patterned pillars on the surface, and therefore it forms a heterogeneously wetted contact surface of alternating liquid/solid contacts and air pockets. As a result the overall solid liquid contact area is significantly reduced [2–5]. In the present work we employ surface patterning to create heterogeneous wetting on metallic surfaces. We show that heterogamous wetting in metallic surfaces leads to reduction in corrosion rates and increase in corrosion resistance.

2 Experimental procedures 2.1 Specimen preparation Pure nickel (99.7 Wt.%) was selected as a model metal. Samples of 1.5 cm × 1.5 cm size were polished to a standard finish with a roughness that did not exceed 50 nm. Laser ablation with a single pulse copper bromide (CuBr) metal vapour laser was used to create special surface textures. During laser ablation, nitrogen (N2) was blown to protect the surfaces from oxidation and clean any debris. The pulse duration was selected at 30 ns. Different hole/pattern sizes were achieved with different laser output power within the range of 20-80 W. Figure 1 shows a schematic of the applied pattern. The diameter of the holes and the inter-hole spacing were varied to obtain different textures. Table 1 presents the selected hole size and the inter-hole spacing. The distances between the holes are labeled L1, L2 and L3, where as the D1, D2, D3 etc are the hole diameters.

Figure 1:

Schematic presentation of the proposed surface texture; D relates to the hole diameter and L defines the inter-hole spacing. Table 1:

Hole diameter (D) µm 5 10 20

The hole sizes and inter-hole spacing. Inter-hole spacing L1, µm 5 10 20

Inter-hole spacing L2, µm 10 20 30


Inter-hole spacing L3, µm 20 30 40


87

For easy identification of the samples the following sample labeling system has been adopted: DxLy, where x is the diameter of the hole in m and the y is the inter-hole spacing in m. 2.2 Corrosion tests The corrosion resistance of the samples was determined using a potentiodynamic polarization corrosion test, wherein the samples were immersed into a 0.5M H2SO4 solution at room temperature (24 ◦C). A conventional three-electrode system was used in which a standard calomel electrode (SCE) served as the reference electrode and a platinum electrode as a cathode electrode. In the tests, the applied potentials were in the range of -0.7 to 1.5 V (with respect to SCE) with a scan rate of 1.0 mV/s. Corrosion rates were calculated in terms of the corrosion current density, Icorr, by using linear polarization resistance curves (LPR). The relationship between Icorr and the polarization resistance, Rp, is obtained from the Stearn–Geary equation [5]:

Rp  where  a and

c

a c E  iapp 2.3( I corr )( a   c )

(1)

are the Tafel slopes of the anodic and cathodic reactions,

respectively, Icorr is the corrosion current density and the

E is polarization iapp

resistance. After the corrosion test the sample surfaces were examined using JEOL 5800 scanning electron microscopy (SEM), and energy-dispersive X-ray spectrometry (EDS).

3 Results and discussion 3.1 Corrosion tests The electrochemical characteristics of the patterned Ni samples were investigated by potentiodynamic techniques described in Sec 2.2. Polarization curves for all patterned samples and the reference sample were analyzed and the corresponding corrosion current density, Icorr, and potential, Ecorr ,were obtained. Figure 2, shows the variation of the Icorr for different hole diameters (1/D) in samples with the same ratio of patterned to non-patterned area (D/L)2. Clearly samples with D=10 µm(1/D=0.1 µm-1) have the worst corrosion properties compared to any other samples with either pattern density((D/L)2=0.25 and (D/L)2=1.0). On the other hand, samples with D=20 µm(1/D=0.05 µm-1) show the lowest Icorr , and therefore better corrosion resistance at either pattern density. In addition, the corrosion resistance observed in samples with D=20 µm exceeds the performance of the reference sample (1/D=0.0 µm-1) which is a significant improvement of the corrosion properties of the metal surface. The observed WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)

88 Materials Characterisation V trends in Figure 2 suggest that a decrease of the hole size (D) µm towards nanoscale size could decrease the corrosion current density, Icorr, even further. Figure 3 presents the Ecorr values for samples with pattern density of (D/L)2=1 and (D/L)2=0.25. The corrosion potential (Ecorr) of all surface patterned samples was lower (less noble) than that of the reference sample (1/D=0.0 µm-1) in all tests.

a)

b)

Figure 2:

The Icorr values versus inverse hole diameters (1/D) for all samples with (D/L)2=1 and (D/L)2=0.25 after the corrosion test.

a)

b)

Figure 3:

The Ecorr values versus inverse hole diameters (1/D) for all samples with (D/L)2=1 and (D/L)2=0.25 after the corrosion test.

3.2 SEM images after corrosion tests The patterned samples were examined by SEM before and after above corrosion tests (see Figs. 4–6). The original REF sample showed some severe, localized corrosion. It is evident that those areas were formed by coalescence of small pits. Sample D10L30 showed a severe corroded appearance (Figs. 5a and 5b). According to Icorr of this sample there is a good agreement between corrosion current density and the surface condition.



a)

89

b)

Figure 4:

a)

SEM images of the reference sample(REF) a) before and b) after corrosion testing.

b)

Figure 5:

a)

SEM images of the D10L30 a) before and b) after corrosion testing.

b)

Figure 6:

SEM images of the D20L40 a) before and b) after corrosion testing.


90 Materials Characterisation V Sample D20L40 exhibited the lowest amount of patterning degradation. Comparing the original sample and after the corrosion test there was no obvious change of the surface pattern [Figures 6(a) and 6(b)]. The measured Icorr for this sample was the lowest of all samples including reference and patterned. No surface damaged was detected also in samples D20L20 and D20L40. Clearly the observed surface damage in all samples was in a good agreement with the predicted trends by the measured Icorr values. Sample REF

a) Sample REF

b) Figure 7:

EDS analysis of the reference sample REF a) before and b) after corrosion testing.

3.3 EDS analysis of the patterned samples In order to investigate the significant improvement of the corrosion resistance in some samples energy-dispersive x-ray spectrometry (EDS) was used. The main goal was to evaluate the change of the O concentration on the surface of the samples and on the inside of the patterned holes. Fig. 7(a) refers to sample REF before the corrosion test. In the spectrum three elements, nickel (Ni), oxygen (O) and carbon (C) are observed. The intensity of the most intense peaks (Ni: Lα=0.851 and O: Kα=0.523 keV) was used to calculate the ratio of Ni/O (see Fig 7(a)). Comparing sample REF before the test [Fig. 7(a)] and after the test [Fig. 7(b)] shows that there is virtually no change in the O concentration on the surface. The WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


91

Ni/O ration remained at 49/5, indicating that passive oxide layer was not formed. Similar EDS spectra were obtained for all corrosion damaged samples. For instance the EDS spectra of sample D5L20 on the surface [Fig. 8 (a)] and in the hole [Fig. 8 (b)] after corrosion testing do not show any significant difference and the Ni/O ratio is the same as for the polished REF sample. However, the EDS spectra for the pattered samples with the improved corrosion properties (D20L20, D20L30, and D20L40) exhibit different behaviour. The Ni/O ratios for sample D20L40 before and after the corrosion test are shown in [Figs. 9 (a),(b) and Figs. 10(a),(b)]. The concentration of the O on the surface before and after the corrosion test slightly changed, which is consistent with the other corrosion damaged samples. However, the bottom of the patterned holes shows significant increase in O concentration (Figs. 8(b) and 10(b)). This clearly indicates that Ni oxides were formed but were not dissolved by the electrolyte. A possible explanation could be that the fluid actually did not reach the bottom of the patterned hole. In other words, in the samples with better corrosion resistance, the contact between the electrolyte and the metal surface is heterogeneous wetting – alternating solid/liquid zones and air pockets.

Sample D5L20-surface

a)

Sample D5L20-hole

b) Figure 8:

EDS analysis of the sample D5L20 after the corrosion test a) surface and b) hole.



Sample D20L40-surface

a)

Sample D20L40-hole

b) Figure 9:

EDS analysis of the virgin sample D20L40 a) surface and b) hole. Sample D20L40-surface

a)

Sample D20L40-hole

b) Figure 10:

EDS analysis of the sample D20L40 after the corrosion test a) surface and b) hole.



93

4 Conclusions Surface patterns with different pattern density were successfully created on Ni by laser ablation. The patterns had a distinct effect on the corrosion properties of the metallic surfaces. A specific group of patterns led to significant decrease of the corrosion current density and the corresponding corrosion rate. For this group of patterns it has been shown that a possible reason for the decrease in corrosion rate is the detected heterogeneous wetting on the patterned surface.

References [1] K. Varanasi, T. Deng, J. David Smith, M. Hsu, and N. Bhate, “Frost formation and ice adhesion on superhydrophobic surfaces” Appl. Phys. Lett. Vol. 97, 2010, No 234102. [2] M. Shafiei and A. T. Alpas, “Nanocrystalline nickel films with lotus leaf texture for superhydrophobic and low friction surfaces”, Applied Surface Science Vol. 256, 2009, pp. 710–719. [3] W. Barthlott and C. Neinhuis, “Purity of the sacred lotus, or escape from contamination in biological surfaces, Planta, Vol. 202, 1997, pp. 1-8. [4] S.J. Lee, C.H. Huang, J.J. Lai and Y.P. Chen, “Corrosion-resistant component for PEM fuel cells”, Journal of Power Sources, Vol. 131, 2004, pp. 162-168. [5] T. Liu, Y. Yin and L. Dong, “New application of the underwater superhydrophobic surface in the corrosion protection”, Advanced Materials Research, Vols. 79-82, 2009, pp. 1115-1118. [6] M. G. Fontana and N. D. Greene, “Corrosion engineering”, 2nd edition, McGraw Hill, 1978.




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Material characterisation to understand various modes of corrosion failures in large military vehicles of historical importance A. Saeed1, Z. Khan1, N. Garland1 & R. Smith2 1

Sustainable Design Research Centre, School of Design, Engineering and Computing, Bournemouth University, UK 2 The Tank Museum, Bovington, UK

Abstract Large military vehicles within museum collections are stored in two distinct environments, controlled and uncontrolled, with an intermittent transitional mode where vehicles travel between the two. Variable environmental conditions combined with operational factors pose significant risks to the reliability, durability and longevity of these vehicles. Although there are methods for retarding or decelerating aspects of failure, to maintain the integrity and originality of these vehicles as artefacts a sustainable methodology for conserving these vehicles should be developed. Corrosion is one of the significant contributors to the structural damage and material aging of historical military vehicles; therefore an experimental study was conducted to understand the prevailing mechanisms of failures due to corrosion with various types occurring in these vehicles identified. This paper presents various modes of corrosion in historic vehicles while X-ray Fluorescence and ultrasonic scanning corrosion mapping techniques characterise corrosion inhibiting materials and subsequent material loss. Understanding material profiles and their link to environmental exposure during use and non use of these vehicles will lead to a sustainable methodology for their conservation. Keywords: sustainable, military, vehicle, corrosion, materials, conservation.



1 Introduction The structural and functional deterioration of the large metal structures due to their environmental exposure and the operating conditions are responsible for their structural aging [1]. Aging mechanisms in large metal structures such as bridges, aircrafts, vehicles etc is recognised as a growing problem [2]. Aging mechanisms such as corrosion, surface and sub surface cracks, wear due to interacting surfaces and undesired stresses are leading to mechanical failures, higher maintenance costs and compromises the structural integrity. Corrosion has been identified as one of the major contributor to the aging mechanisms; it affects all metal structures by deteriorating the material’s properties [3]. This research focuses upon the corrosion damage within the large historic vehicles in the museum environment. Corrosion is indiscriminate in affecting metal structures and museum artefacts are no exception. The military vehicles from World War I and II were designed and manufactured according to the automotive technologies in place at those times [4] with only a limited lifeexpectancy, however their participation in the war and the subsequent historic significance means it is important to devise sustainable methodologies for longterm preservation. Continuous corrosion problems within large historic military vehicles in the museum environment are leading to the concerns of their longevity, in particular where failures in protective measures are also identified [5]. For the purpose of sustainable methods it is important to understand the various corrosion modes, material characterisation and the material loss profile due to corrosion. For this research military vehicles were inspected for visual signs of corrosion and novel methods such as X-ray fluorescence for elemental identification and ultrasonic scanning for material loss were conducted. Visual inspection identified different modes of corrosion depending on their exposure to extreme environments during the war, operating conditions at the time of their service life, types of materials used and their current working or non-working environments. X-Ray fluorescence method found Fe, the main constituent; other elements such as W, Mo and Mn were also identified. Ultrasonic methods found severe material loss due to corrosion in the selected samples.

2 Experimental methodologies The following methods are used in the research. Only the most common and reoccurring types of corrosion which were found on a larger scale in the vehicles are presented here. 2.1 Visual inspection Visual inspection was completed on more than 20 large military vehicles from the World War I, II and post war era. The investigation found one or more forms of corrosion in the same vehicles. Some widespread and most occurring modes of corrosion are described from four military tanks Table 1. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


Table 1:

97

Military tanks presented in this research.

No

Vehicle

Environment

1 2 3 4 5

Mark II Sherman* King Tiger Centurion Scorpion

Controlled Uncontrolled Controlled Uncontrolled Uncontrolled

Manufacture

Country of Origin

Era

1917 United Kingdom WW 1 1940-1941 United States WW 2 1943 Germany WW 2 1945 United Kingdom Post war 1973 United Kingdom Post War

XRF and Ultrasonic scanning is conducted on Sherman.

2.2 X-ray fluorescence X-ray fluorescence was conducted on Sherman M4A1 for elemental identification. The test sample was bombarded by an intense x-ray beam resulting in the emission of fluorescent x-rays from the test object. The emitted fluorescent x-rays are then detected by XRF analyser [6] which identifies the elements by measuring the energies of the emitted fluorescent x-rays and counting the number of rays in each energy spectrum [6, 7]. 2.3 Ultrasonic scanning Ultrasonic scanning uses pulse waves from 0.10 to 15.0 MHz frequencies range to measure material loss and other dimensional anomalies in the test objects [8]. This research work presents ultrasonic scanning using pulse echo method where pulses are transmitted and received back on the same side of test object to detect and characterise defects.

3 Results 3.1 Uniform corrosion With uniform corrosion a direct chemical attack affects the surface of the metal and then spreads out evenly across the entire surface, results are not usually fatal and it is not classified as one of the severe forms [9]. Figure 1 of the King Tiger’s flywheel shows uniform corrosion spreading across the surface. Uniform corrosion was also found in Tog II and FT 17 on considerably larger surface areas. Uniform corrosion, if left to continue, could lead to general thinning of the surface and may transform in to another form of corrosion [10]. 3.2 Fretting corrosion Fretting corrosion occurs at the interface of the interacting surfaces which are subject to relative motions. The King Tiger planetary gears (Figure 2) are a classic example of this form. Interacting mechanical wear lead to the removal of material particles, followed by the oxidation of the material debris and the newly exposed surface [11]. The oxidised debris would act as an abrasive media leading to further wear and fatigue failures could occur [9, 12]. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


Figure 1:

Figure 2:

King tiger flywheel.

King tiger planetary gears.

3.3 Stress corrosion cracking Stress corrosion cracking is considered one of the severe types of corrosion, it could cause a partial or complete failure of a structure especially in the presence of dynamic or static loading and initiation of a stress corrosion crack [13]. The Mark-II Tank glacis plate in Figure 3 shows stress corrosion cracking where a crack is initiated and penetrating through the surface with no significant material loss on the surface. Stress corrosion is difficult to detect, hard to predict and unexpected failure can occur [13, 14]. 3.4 Inter-granular/exfoliation corrosion Inter-granular corrosion describes the attack on a metal’s or alloy’s grains or boundaries [15]. Corrosion products between the grain boundaries exert pressure on the grains and the result is exfoliation corrosion. The Centurion armoured skirt in Figure 4 is an example of exfoliation corrosion leading to lifting or leafing effect. Leafing or lifting effect is a direct result of the expanding corrosive products among the grains of the metals which forces the grains apart [9]. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


Figure 3:

Figure 4:

99

Mark II glacis plate.

Centurion armoured skirt.

3.5 Galvanic corrosion Galvanic corrosion is the electrochemical reaction between two dissimilar metals in the presence of electron conductive path and the preferential corrosion of one metal more than the other [16]. This mode of corrosion was identified in the Scorpion, see Figure 5. The presence of moisture among the bolt, washer and the armoured skirt has caused this type of corrosion. 3.6 Pitting corrosion Centurion in Figure 6 shows signs of pitting corrosion. In this case the corrosion is confined to a small point with resulting cavities in the metal. The localised phenomenon of corrosion creates holes in the metal and then penetrates inwards, deteriorating metal from the inside. The corrosion products covers the pit and creating a small hole with no significant material loss on the surface leading to a catastrophic malfunction, it is one of the severe forms of corrosion which is very difficult to detect and hard to predict the result [14, 17, 18].



Figure 5:

Galvanic corrosion in scorpion.

Figure 6:

Pitting in centurion.

3.7 X-ray Fluorescence of the Sherman M4A1 X-Ray fluoresce (XRF) was conducted on three points A, B, C on corroded, E, F, and G on sand blasted surfaces and point D on the cross section of the sample, see Figure 7. Table 2 shows results of the points A, B and C. On averages Fe 92.07% is the highest, second highest is the Si 5.39%, Mn 0.91% was the third highest and the lowest was Cu with a trace of 0.04%. Figure 8 shows the graphical representation of the XRF conducted at point D where Fe 98.30% is the highest, second highest is Si 0.65%, Mn 0.63% the third highest and traces of Ti, V, Co and Mo were identified. XRF conducted at points E, F and G on the sandblasted surface shown on averages Fe 98.12% the highest, Mn 0.66% second highest, Si 0.59% the third highest while W 0.02% the lowest and no traces of Co or V were identified. Figure 9 shows the graphical representation of the XRF analysis at point G.



Figure 7: Table 2:

101

XRF points on Sherman.

XRF of the corroded surface of M4A1.

Vehicle Type: Sherman Sample- 1 Time duration: 30 seconds

Surface: Corroded MQ for points: A= 5.70 B= 4.50 C= 4.60 06 American 06 American

Material Standard

06 American

Constituents %

A- Reading 254

B-Reading 255

C-Reading 256

4.89 0.46 0.19 0.05 0.11

5.92 0.54 0.19 0.05 0.08

5.38 0.85 0.2 0.05 0.09

1 2 3 4 5

Si P Ti V Cr

6

Mn

0.86

1.11

0.78

7 8 9 10

Fe Co Ni Cu

92.68 0.21 0.06 0.05

91.48 0.22 0.08 0.03

92.05 0.12 0.1 0.04

11 12

Mo W

0.05 0.33

0.06 0.2

0.06 0.12

3.8 Pulse echo ultrasonic scanning Pulse Echo scanning was conducted on the Sherman M4A1 with a probe of 4.00 MHz twin crystal and 10.00 mm diameter using 0.25 MHz Hi pass filter and 10.00 MHz low pass filter frequencies with a probe delay of 7.630 µS. Compression waves with a velocity of 5903.00 M/sec are used to scan a total area of 174.00 mm length with 6.040 pulses/mm and 68.00 mm width with 3.580 pulses/mm mapping material loss every 2.00 mm on the sample. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


Figure 8:

XRF analysis of point D, No Fe (98.30%) included.

Figure 9:

XRF analysis of point G, No Fe (98.09%) included.

Figure 10 shows the end views and the c-scan of the sample with implementation of colour coding Red for 0.00 mm and Blue for 10.00 mm. Two cursors Red and Blue show the X-axis, Y-axis location and the corresponding depths of the particular points on the sample. A user defined area is taken for finding the maximum and minimum thicknesses starting at 2.00 mm on the X-axis, 2.00 mm on the Y-axis in Figure 11 and finishing at 172.00 mm on the x-axis and 66.00 mm at the y-axis in Figure 12 respectively to avoid any distortion in signals at the edges. The scan detected maximum remaining thickness of 9.80 mm at horizontal position of 150.00 mm and vertical position of 26.00 mm with a drastic loss in material to a minimum of 4.60 mm at horizontal position of 28.00 mm and vertical position of 8.00 mm. The maximum thickness of 9.80 mm is recorded at two different points. The second lowest depth is 4.70 mm and is detected at seven different WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


Figure 10:

Figure 11:

103

Pulse echo scan of M4A1.

User defined area starting at 2.00 mm on x-axis and 2.00 at y-axis.

positions; third lowest of 4.80 mm at six various points and the remaining thickness of 4.90 mm are recorded at four different points in the samples. The minimum remaining thickness of 4.60 mm is recorded at three different points.

4 Discussion The modes of corrosion identified above can be retarded by using the protective measures such as coatings and paint for uniform corrosion, application of lubrication and exclusion of air in the fretting corrosion, alleviating the stress in stress corrosion cracking, using same metal types in galvanic or applying insulation between two different metals. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


Figure 12:

User defined area finishing at 174 mm on x-axis and 66 at y-axis.

The alloying elements such as Mn, Cr, V W and Mo identified though XRF are used as hardening agents to improve hardness, ductility and tensile strength [19]. Also Mo, Mn and P are used for the rust and corrosion prevention. Mo is used in high strength alloys and super alloys for strength and wear resistance [20]. Sherman M4A1 was designed by Lima Locomotive Works in the United States. At the time of manufacture the armoured thickness range was kept between maximum of 62.00 mm and a minimum of 12.00 mm [21, 22]. However it is not practical to follow any possible modification in the armour thickness and because of this reason the highest thickness recorded in the scan is considered the maximum thickness. During the service life, exposure to extreme environments has led to the deterioration and material loss. The ultrasonic scan of has shown a remarkable material loss of 5.20 mm between the maximum remaining thickness of 9.80 mm and the minimum remaining thickness of 4.60 mm. Currently in the museum collection, the Sherman travels between two controlled and uncontrolled environments. Varying environments and the likelihood of failing protective measures, temperature fluctuations, humidity ratio variations, rain water and the extent of chemical reaction etc will influence the rate of corrosion and in the controlled environment corrosion residues from may result in indoor atmospheric corrosion [23, 24].

5 Conclusion The research work uses a conventional method of analysing corrosion through surveying vehicles in the museum environment which is followed by novel methods of XRF for the elemental identification and Ultrasonic scanning to measure and map corrosion. Research work will be conducted for corrosion WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


105

modelling and for the life expectancy of large historic vehicles based upon the results from this research work in order to maximise the longevity, durability and reliability.

Acknowledgements The authors would like to thank Mark Nel, Mark Clark in AGR, Jon Previtt West-Dean College and Mike Hayton in The Tank Museum Bovington for their in kind support during this research work.

References [1] M. Colavita, "Service Life and Aging of Military Equipment," Corrosion: Environments and Industries (ASM Handbook), vol. 13C, pp. 220-228, 2006. [2] A. S. Vinod, "Corrosion in the Military," presented at the Corrosion: Environments and Industries (ASM International), 2006. [3] Z. Ahmad, "Basic Concepts in Corrosion," in Principles of Corrosion Engineering and Corrosion Control, Oxford: Butterworth-Heinemann, 2006, pp. 9-56. [4] I. C. Handsy and J. Repp, "Ground Vehicle Corrosion," Corrosion: Environments and Industries (ASM Handbook), vol. 13C, pp. 148-150, 2006. [5] M. T. Gudze and R. E. Melchers, "Operational based corrosion analysis in naval ships," Corrosion Science, vol. 50, pp. 3296-3307, 2008. [6] B. Beckhoff, Handbook of practical X-ray fluorescence analysis. Berlin; London: Springer, 2006. [7] R. v. Grieken and A. A. Markowicz, Handbook of X-ray spectrometry, 2nd ed., rev. and expanded. ed. New York, N.Y.: M. Dekker, 2002. [8] J. C. Drury, Ultrasonic flaw detection for technicians, 3rd ed. [Swansea]: Silverwing Ltd., 2004. [9] R. Baboian, Corrosion tests and standards : application and interpretation. Philadelphia, Pa.: ASTM, 1995. [10] H. H. Uhlig and R. W. Revie, Corrosion and corrosion control : an introduction to corrosion science and engineering, 4th ed. / R. Winston Revie. ed. Hoboken, N.J.: Wiley-Interscience; Chichester: John Wiley [distributor], 2008. [11] R. C. Bill, "The role of oxidation in the fretting wear process," 1981. [12] A. Neyman and O. Olszewski, "Research on fretting wear dependence of hardness ratio and friction coefficient of fretted couple," Wear, vol. 162164, pp. 939-943, 1993. [13] R. P. M. Procter, et al., "Stress-corrosion cracking of C---Mn steels in methanol-ammonia environments--I. Effects of environmental and mechanical variables," Corrosion Science, vol. 33, pp. 1009-1031, 1992. [14] R. C. Newman, "Stress Corrosion Cracking," in Shreir's Corrosion, J. A. R. Tony, Ed., Oxford: Elsevier, 2010, pp. 864-901. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)

106 Materials Characterisation V [15] C. A. e. Hampel, The encyclopedia of electrochemistry. NY: [s.n.], 1964. [16] H. P. Hack, "Galvanic Corrosion," in Shreir's Corrosion, J. A. R. Tony, Ed., Oxford: Elsevier, 2010, pp. 828-856. [17] Z. Szklarska-Smialowska, Pitting corrosion of metals. Houston, Tex.: National Association of Corrosion Engineers, 1986. [18] R. Baboian, "Automotive corrosion tests and standards," 1996. [19] M. F. Ashby and D. R. H. Jones, Engineering materials 2 : an introduction to microstructures, processing and design, 3rd ed. Oxford: ButterworthHeinemann, 2006. [20] F. F. Schmidt, The engineering properties of molybdenum and molybdenum alloys. Columbus, Ohio: Defense Metals Information Center, 1963. [21] P. Chamberlain and C. Ellis, British and American tanks of World War II: the complete illustrated history of British, American, and Commonwealth tanks, 1939-1945. London New York: Cassell & Co. ;Distributed in the USA by Sterling Pub., 2000. [22] R. Jackson, Tanks : and armoured fighting vehicles. Bath: Parragon, 2007. [23] A. R. Mendoza and F. Corvo, "Outdoor and indoor atmospheric corrosion of non-ferrous metals," Corrosion Science, vol. 42, pp. 1123-1147, 2000. [24] I. Odnevall and C. Leygraf, "The formation of Zn4Cl2(OH)4SO4 · 5H2O in an urban and an industrial atmosphere," Corrosion Science, vol. 36, pp. 1551-1559, 1994.


Section 4 Computational models and experiments



109

A multi-factor interaction model (MFIM) for damage initiation and progression C. C. Chamis NASA Glenn Research Center Cleveland, USA

Abstract A Multi-Factor-Interaction-Model (MFIM) is briefly described to represent complex point material behavior in a single equation. The model is of product form in order to represent coupled interactions and to be computationally effective. The model describes a continuum or surface in space that represents the complex material behavior in terms of the various factors that affect a specified material behavior. The material specified behavior is inclusive of all material properties, mechanical, thermal, physical and effects thereon, such as temperature, time, cyclic loadings, etc. Sample case results simulated by using MFIM are compared with test data to illustrate its versatility and its relevance to reality. These results show that the MFIM can accurately predict metal matrix composite fatigue data and mechanical properties of a steel alloy. Helpful guidelines for its effective use are also included. Keywords: material properties, high temperature, nonlinearities.

1 Introduction The simulation of complex material behavior resulting from the interaction of several factors (such as temperature, nonlinear material due to high stress, time dependence, fatigue, etc), has been mainly performed by factor-specific representations. For example, entire text books are devoted to plasticity, creep, fatigue and high strain rate to mention only a few. Investigators have derived equations that describe material behavior for each factor-specific effect. Suppose we visualize that the material behavior is a continuum represented by some surface. Then, we can think of some representation which describes that surface which is inclusive of all participating factors that affect material behavior either singly or interactively in various combinations. To that end, research has WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line) doi:10.2495/MC110101

110 Materials Characterisation V been a continuing activity at Glenn Research Center for about twenty-five years. It started with a primitive form of MFIM representation for describing complex composite behavior in polymer matrix composites [1]. It was extended to metal matrix composites [2] and continues to be evolved in Aerospace Plane and the High Speed support of the National Research Programs. The result of all this research is that general guidelines for its usage are briefly described. Simulation results are presented and compared (where available) with experimental data to illustrate its versatility and perhaps demonstrate the claim for its uniqueness.

2 Fundamental considerations and the (MFIM) We start with the premise that if we are to quantify the range of factors affecting material properties, we need a description of material behavior. In this context, it is reasonable to consider that material behavior constitutes an n-dimensional space (Material Behavior Space (MBS)) where each point represents a specific aspect of material behavior. It is further reasonable to assume that MBS can be described by an assumed interpolation function. One convenient interpolation function is a polynomial of product form because mutual interactions among different factors can be represented by the overall product, and includes those cross products in common algebraic polynomials. In this investigation, MBS is assumed to be described by the following multifactor interaction equation (MFIM): N MP   Aimi (1) M P0 i 1 where M P is the property affected to be evaluated. MPo corresponds to the initial (reference) material state or condition. Ai represents the ith factor that influences material behavior, and mi is an exponent. Ai is further defined by:  B (2) Ai   1    B0  Here B represents a specific cause factor for behavior (for example, temperature), and Bo is the corresponding final value. This concept is schematically represented in Fig. 1, the development of the multi-factorinteraction model (MFIM) to represent complex material behavior by a single equation. The objective of the present paper is to briefly describe MFIM and present results obtained there from to illustrate its uniqueness and its versatility. Values for Bo and mi for specific behavior are selected either from known behavior or more likely from a best judgment in conjunction with consultations with seasoned professionals for that behavior. By representing the MBS with the MFIM of product form (Eq. (1)), we gain another distinct advantage. The behavior factors, B, can also be represented by another level of MFIM or progressive substructuring of equation (1). The progressive substructuring leads to a multitier representation of the MBS that permits intrinsic lower tier behaviors to influence more than one factor at the next higher tier. In other words, the observed specific behavior (Bi may depend on another set of lower WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


111

tier elemental behaviors). Further, the behavior factors in this lower set of specific behaviors may depend on yet another next lower tier of elemental behaviors. That is, there are usually sets and subsets of specific behaviors that hierarchically influence the higher level behaviors. When this is done, N can be limited to 6, (for example), but the number of factors influencing material behavior at the next lower tier will increase exponentially as Nj where j is the number of 6-factor tiers. For example, when j = 3, N = 216, and so forth. This representation is natural for multiparallel processing computers where the tiers are programmed with different granularities. Obviously, then, the motivation for selecting such a form is for computational and programming effectiveness. Another reason for selecting an MFIM of product form is that the effect of each factor can be evaluated separately. The interpretation of Bo is that it represents a scale, whereas mi represents a shape or path. For example, (1 – B/Bo)mi where 1 > B/Bo and + ∞ < mi < – ∞, covers the whole space as is illustrated in Fig. 2.

Figure 1:

Conceptual schematic of material behaviour through a MultiFactor Interaction Model (MFIM).

The inclusiveness of this particular form, combined with its simplicity, makes it very attractive for a computational simulation. An expanded form of MFIM is shown in Eq. (3) below: m

Mp  Tgw  T    Mpo  Tgw  To   E ...1  e  E ef 

u

  

  1  Sf  Ce 1   Cef

n

  

p

 t 1   Sf tf

  

v

  ... 

  

 Cc 1   Ccf

  M NM 1   Sf Nf M

q

  

r

 T NT  1    Sf Nf T 

s

  1   f

  ... 

The Multifactor Equation

(3)

w


112 Materials Characterisation V for inclusiveness and convenience of presentation. The top line includes six factors in parentheses: ( Tgw – T) / (Tgw – T0 ) denotes temperature effects: (1 – σ/Sf ) denotes combined stress effect; (1 – σ t / Sf tf ) denotes time dependent effects: ( 1 – σ m Nm / Sf Nfm ) denotes mechanical cyclic load effects where σm is the cyclic stress due to NM cycling; (1 – σT NT / Sf Nft ) denotes thermal stress due to NT cycling; ( 1 – ω / ωf ) denotes frequency effects due to ω – frequency. The factors in the second line denote erosion, corrosion and chemical or metallurgical effects. The dots between the first and second lines indicate that several other factors can be included. Suffice it to say that the MFIM is generic and inclusive. Two points to be noted are: (1) not all terms have to be included; and (2) substructuring may be appropriate after six factors. For example, the three factors in the second line affect some, if not all the factors in the first line. This definitely is the case for the chemical/metallurgical factor.

Figure 2:

Effect of variation in the exponent.

3 Guidelines for usage of the (MFIM) Some general guidelines for usage of the MFIM are appropriate: (1) Several factors may be programmed for inclusiveness. (2) Factors that do not contribute to that simulation assign zero (0) exponents. (3) Selecting exponents may be intriguing in the absence of data. Start with some guess of expected behavior and let the feedback from the material property behavior to guide the next update. (4) The exponents can also be evaluated if there is some data available especially from combined testing. This can be done by expressing the expanded equation using logs and then evaluate them by using the least squares method WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


113

since the exponents will be the coefficients of linear algebraic equations. (5) To simulate block fatigue evaluate the degradation up to the end of the first block then use the degraded properties for the second block and repeat for additional blocks – degradation after second input to third etc. (6) Unloading generally requires different exponent then loading. Select the unloading exponent to simulate the expected loading-unloading by stress cycle. (7) Intermittent time effects are handled the same way as block fatigue. (8) Combined fatigue (thermo mechanical) is simulated by using the corresponding cycles and stresses in the respective factors. (9) Frequency effects are simulated in combination with cyclic loads. Stress value is not needed here since it is included in the cyclic effects factor. (10) The static stress and the temperature factors when present are used in combination with all active factors for that simulation effect. (11) To simulate fabrication process use the temperature factor, the stress factor due to pressure, a flow model and update the geometry through the finite element model of the component. The user will invent other ways to use MFIM that the author has not even thought of yet. The most important point to remember is that MFIM is to be used incrementally where the current values are updated from global thermo structural, etc. analyses. This may require a two level iteration: (a) local to achieve local equilibrium. During the iteration, the factors are changed until local equilibrium is reached; and (b) global to achieve global equilibrium with respect to boundary loads and supports conditions and internal stress field or energy input in the increment equals the energy added to the structure. Another important point to note is that the exponent for specific factor in MFIM represents a general trend and not the entire precise path from its reference value to its final value. Also, note that the MFIM is used for all the properties that is the same factors but with different exponent. Results from sample cases for MFIM application are described for: metal matrix composites (MMC), and a nickel base alloy. As mentioned previously, these results are presented to illustrate a few of the generic features of the MFIM applicability to convey to the reader what has be done and infer therefrom what can be done. The specific how is described in the references.

4 MFIM application to MMC In this sample case MFIM is part of the METCAN (Metal Matrix Composite Analyzer) [3]. The simulation for the properties to be shown starts from the fabrication process as depicted in Fig. 3. The constituent material properties used are summarized in Table 1. MMC Simulation of the transverse strength is shown in Fig. 4. For this simulation the factors for temperature, stress, and time were activated with respective exponents about 0.5. It can be observed that the simulation predictions track the data almost exactly. That simulation is very complex because it incorporates matrix, interphase and fibers. It starts from the fabrication process and ends with monotonic loading to fracture. Sample simulation results for fatigue of the same composite are compared with data in Fig. 5. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


Figure 3: Table 1:

METCAN Metal Matrix Composite Behavior – computational simulation sequence. Constituent (fiber/matrix) material properties used in METCAN.

It is observed that the agreement is very good. In addition, the simulation identifies the types of failures that lead to specimen fracture that is achievable by the use of the MFIM. For this case, the mechanical cycle term in the MFIM was also activated with an exponent of about 0.5 for all the constituent properties. Simulation results for stress at times are compared in Fig. 6. As can be seen the agreement is very good. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


115

Figure 4:

Transverse stress strain curve of SiC/Ti-6-4 T = 73 oF, FVR = 0.34.

Figure 5:

Predicted isothermal fatigue life [0] SCS6/Ti-24A1-11 Nb, 0.35 FVR (70 oF (23 oC), R = 0.1).

Figure 6:

Creep behavior of [0] SCS6/Ti-24Al-11Nb composite at 815oC and 310 MPa; FVR = 0.35; stress-free temperature = 815oC.



Figure 7:

Calibration of nickel base alloy 713C elastic modulus as a function of temperature.

Figure 8:

Calibration of nickel base alloy IN-100 coefficient of thermal expansion as a function of temperature.

The time factor was activated for the fabrication process and the tensile load to fracture. Additional simulation results are described in [3, 4].



117

5 Properties for nickel base alloy Simulation results of temperature effects on nickel base alloy 713C properties [5] are shown respectively in Fig. 7 for modulus, Fig. 8 for thermal expansion coefficient, Fig. 9 for thermal heat conductivity, and in Fig. 10 for heat capacity. Collectively the comparisons are in good agreement. Collectively these results show the effectiveness of MFIM.

Figure 9:

Calibration of nickel base alloy MAR-M 200 thermal conductivity as a function of temperature.

Figure 10: Calibration of nickel base alloy IN-100 specific heat as a function of temperature. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


6 Concluding remarks The salient concluding remarks from a description of the multi-factor interaction model (MFIM) to predict complex material behavior by using a single equation and select applications are: (1) MFIM is a material point simulator. It is unique and inclusive for both composites and metals. (2) It is used incrementally in combination with analyses that provide local stress information. (3) It usually requires alternative solutions consisting of both local and global convergence. (4) Applications for meta matrix composites under different loading conditions predicted results that were in very good agreement with test data. (5) Familiarity and confidence and extensive usage will be gained by continuing usage application to complex solutions of material behavior and to simulate difficult and even untried problems.

References [1] Chamis, C.C., Lark, R.F. and Sinclair, J.H.: Integrated theory for Predicting the Hydrothermal Mechanical Behavior of Composite Structural Components. ASTM pp. 160-192. [2] Chamis, C.C. and Hopkins, D.A.: Thermo Viscoplastic Nonlinear Constitutive Relationships Structural Analysis of High Temperature Metal Matrix Component. NASA TM 87291, 1985 [3] Chamis, C.C., Murthy, P.L.N. and Hopkins, D.A.: Computational Simulation of High Temperature Metal Matrix Composites Cyclic Behavior. ASTM, STP 1080, 1990, pp 56–69. [4] Tong, M.T., Singhal, S.N., Chamis, C.C. and Murthy, P.L.N.: Simulation of Fatigue Behavior of High Temperature Metal Matrix Composites. ASTM Reprint from Standard Technical Publication, 1253, 1996, pp. 540-551. [5] Boyce, L. and Chamis, C.C.: Probabilistic Constitutive Relationships for Cyclic Material Strength Models. AIAA.ASME/ASCE/AHS 29th Structures, Structural Dynamics and Materials Conference. Part 3, AIAA, 1988, pp. 1299-1306. [6] Progressive Fracture Structural Analysis of National Wind Tunnel Structures by L. Minnetyan, NASA CR 198485, May 1996.



119

Analytical solution of a two-dimensional elastostatic problem of functionally graded materials via the Airy stress function H. Sakurai Co-operative Education Centre, Sendai National College of Technology, Japan

Abstract Functionally Graded Materials (FGMs) possess properties that vary gradually as a function of spatial coordinates. They are different from conventional composite materials in that they have no distinct interfaces at which their material properties change abruptly. These FGMs are suitable for various applications, such as aerospace, nuclear fusion, biomaterial electronics, etc. In practice, applications of analytical solutions are limited. However, the analytical solutions are very important as standards for evaluating numerical simulation results and they are also important to mathematical understanding. Little research on the analytical solutions of two-dimensional elastostatic problems has been reported. Furthermore, few analytical solutions using Airy stress functions have been published. The purpose of this paper is to propose an analytical method for the two-dimensional elastostatic problems of FGMs using the Airy stress function. In the present investigation, FGMs in which the properties of the materials vary exponentially in one direction are examined. A few numerical examples are presented and the validity of the method is shown by comparisons with the results of past studies. Keywords: analytical solution, functionally graded material, two-dimensional problem, Airy stress function.

1 Introduction The Functionally Graded Materials (FGMs) possess properties that vary gradually as a function of spatial coordinates. They are different from conventional composite materials in that they have no distinct interfaces at which WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line) doi:10.2495/MC110111

120 Materials Characterisation V their material properties change abruptly [1]. These FGMs are suitable for various applications, such as aerospace, nuclear fusion, biomaterial electronics, etc. Chakraborty et al. developed a new beam element solving FGM beam structures based on the first order shear deformation theory [2]. Nguyen et al. proposed a first-order shear deformation plates models for modeling structures made of FGMs [3]. Xinag and Yang studied free and forced vibrations of a laminated FGM Timoshenko beam with variable thickness under heat conduction [4]. Sankar and his co-workers have been reporting analytical methods for the thermomechanical analysis of FGM beams [5–7]. Zhu and Sankar presented an elasticity solution of a simply supported FGM beam having variation of Young’s modulus distributed by a polynomial in the thickness direction [6]. In Ref. [6], the Fourier series method is used to reduce the governing partial differential equations to the ordinary equations that are then solved by the Galerkin method. Miers and Telles proposed the Boundary Element-Free Method belonging to a meshless technique, for two-dimensional elastostatic analysis of FGMs [8]. Zhong and Yu presented explicit solutions of a cantilever FGM beam having arbitrary graded variations of material properties distributed in the thickness direction based on two-dimensional theory of elasticity [9]. Applications of analytical solutions are limited to practical shapes of analysis regions and boundary conditions. However, the analytical solutions are very important as standards for evaluating numerical simulation results, such as finite element method etc., and they are also important to mathematical understanding. Little research on the analytical solutions of two-dimensional elastostatic problems has been reported. Furthermore, to the best of the author’s knowledge, few analytical solutions using Airy stress functions have been published. The objective of the present paper is to describe the analysis of the twodimensional elastostatic problems of FGMs using the Airy stress function. In this study, FGMs in which the properties of the materials change exponentially in one direction are treated. A few numerical examples are presented and the validity of the method is shown by the comparisons with the results of the present method and results of past studies.

2 Two-dimensional problem of FGMs and basic equations In the Cartesian coordinate system O-xz , we consider a simply supported FGM beam subjected to a transverse load as shown in Fig. 1. The length in the x direction is l and the length in the z direction is h . In Fig. 1, the uniform transverse load and the simply supported boundary condition are one of the examples. In the absence of body forces the equilibrium equations are given as

 xz  zz  0 x z where  xx ,  zz ,  zx are stress components and  zx   xz .  xx  zx 0  z x


(1)


121

p z ( x)  p 0

O

x

h l

z Figure 1:

An example of a simply supported FGM beam.

The relationships between strains and displacements are

 xx  where

u x

 xx ,  zz ,  zx are

 zz 

w z

 zx 

strain components and

u w  z x

(2)

u, w are the displacement

components in the x and the z direction respectively. The strain components should also satisfy the following compatibility condition.

 2  xx  2  zz  2 zx   0 zx z 2 x 2

(3)

The constitutive equations are given as

 xx  s11 xx  s13 zz where

 zz  s13 xx  s33 zz

 zx  s44 zx

(4)

s11 , s13 , s33 , s 44 are elastic moduli. The material properties of FGMs

change gradually as a function of spatial coordinates. We assume material properties varying exponentially in the z direction, i.e. z

s ij  s  F ( z )  s  e h

(5)

where sij are their corresponding values in the plane

z  z 0 with F(z0 )=1 and

0 ij

0 ij

0

F(z) , is called graded function, which expresses the distribution of material properties, and the parameter α is called the graded index [9]. Now, we introduce Airy stress function Φ ( x,z ) expressed by the following equations.

 xx

 2  2 z

 zz

 2  x 2

 zx


 2  zx

(6)

122 Materials Characterisation V Substituting Eqns. (4)-(6) into Eqn. (3), the next governing equation with respect to Airy stress function  ( x , z ) is obtained. 4 3  4  4 0 0 0   0    s s s s     ( 2 ) 2 13 44 33 11 h z 3 z 4 x 2 z 2 x 4 2 2   3  2  2 0 0  0  s s  (2 s130  s 44 )   0 11 13 h x 2 z h 2 z 2 h 2 x 2

s110

(7)

The above equation is the same form as the governing equation of the plates bending with variable rigidity, and it is possible to adopt the same method for the solution [10].

3 Method of solution In Eqn. (7), we introduce the following Airy stress function Φ ( x,z ) . 

 ( x, z )   Ai ( z )  f i ( x)

(8)

i 1

The function f i ( x) should be assumed to satisfy the mechanical boundary conditions at the x  0 and x=l . For instance, under the simply supported condition and the cantilever condition, it should be chosen to satisfy σxx=0, σzz=0 at the both ends and σxx=0 and σxz=0 at the free end respectively. 3.1 Formulation for simply supported FGM beam subjected to uniform pressure In this problem, we assume the following function 

   Ai ( z )  f si ( x )  i 1

where

i 



 A ( z )  sin  x

i 1, 3, 5

i

i

(9)

i , f si ( x)  sin  i x . And the stress components are as follows l  d 2 Ai ( z )  xx    f si ( x) dz 2 i 1, 3, 5

 zz 





i 1, 3, 5

 zx 





i 1, 3, 5

2 i

Ai ( z )  f si ( x )

 i

dAi f ci ( x) dz


(10)


123

f ci ( x)  cos i x . Substituting Eqn. (9) into Eqn. (7), the following ordinary differential equation with respect to the unknown function Ai (z ) is where

obtained easily. 3 2 2 d 4 Ai 2 d Ai 0  d Ai 0  0 0 s s s s  2  {  ( 2  )  } 11 11 13 44 i h dz 3 dz 4 h2 dz 2 2 2  dAi 0 0  (2 s130  s 44 ) i  ( s33  i 2  s130 2 ) i 2 Ai  0 h dz h

s110

(11)

The solution of the above equation is given as

Ai ( z )  C i ,1e i ,1  C i , 2 e i , 2  C i ,3 e i , 3  C i , 4 e i , 4 r z

where the

r z

r z

r z

(12)

ri , j ( j  1, 2, 3, 4) are the roots of the following the 4-th order

equation and the

C i , j ( j  1, 2, 3, 4) are arbitrary constants to be determined by

the boundary conditions at z   h 2 and z  h 2 .

 2  2 1 3 0 ) i2 ri , j ri , j   2  0 (2 s130  s 44 h s11 h  1  1 0 2 2 0  0 (2 s130  s 44 ) i2 ri , j  0 ( s 33  i  s130 2 ) i2  0 h s11 s11 h

ri , j  2 4



Considering Fourier expansion of the transverse pressure

(13)

p z ( x)  p0 , the

following equation is obtained.

pz 

4 p0  1 1 f si ( x)   sin  i x  i 1,3,5 i  i 1,3,5 i

4 p0





(14)

The boundary conditions at the both surfaces are as follows.  zz   p0 ,  zx  0 at z   h 2

 zz  0

,

 zx  0

at z  h 2

(15)

Substituting Eqn. (10) and Eqn. (14) into Eqn. (15), the following four equations for determining the arbitrary constants C i , j are derived.

4 p0 h Ai ( )  2  i 2 i h Ai ( )  0 2

dAi h ( )  0 dz 2 dAi h ( )0 dz 2


(16)


3.2 Formulation for cantilever FGM beam subjected to uniform pressure We deal with the formulation of a cantilever FGM beam such as shown in Fig. 2.

p z ( x)  p0

O

x

h l

z Figure 2:

An example of a cantilever FGM beam.

It is necessary to assume the function f i ( x) satisfying the mechanical boundary conditions at the free end. One of the simplest functions satisfying the above conditions is

  A2 ( z )  f 2 ( x)  A2 ( z )  x 2 where

(17)

f 2 ( x)  x 2 . The stress components are as follows.

 xx   zz   zx  

d 2 A2 2  2 d 2 A2   f ( x )  x 2 z 2 dz 2 dz 2

d 2 f 2 ( x)  2  A ( z )   A2 ( z )  2 2 x 2 dx 2

dA ( z ) df 2 ( x) dA ( z )  2  2   2  2x zx dz dx dz

(18)

Substituting Eqn. (17) into Eqn. (7), the following ordinary differential equation is derived again. 3 2 2 d 4 A2 0  2 d A2 0 0 0  2 d A2 s x  2s11 x  {2(2s13  s44 )  s11 2 x } 2 h dz 4 dz 3 h dz (19) 2  0  dA2  2(2s130  s44 )  2s130 2 A2  0 h dz h 0 2 11



125

We consider the above equation as the following ordinary differential equation with respect to z at the given section x  x0 (constant). 3 2 2 d 4 A2 2 d A2 2 d A2 0 0 0  0  s x  2s11 x0  {2(2s13  s44 )  s11 2 x0 } 2 h dz 4 dz3 h dz (20) 2  0  dA2  2(2s130  s44 )  2s130 2 A2  0 h dz h 2 0 11 0

The solution of the ordinary differential equation is given in the same form as Eqn. (12). The boundary conditions on the top z   h 2 and on the bottom

z  h 2 at the x  x0 (constant) are expressed as follows.

h A2 ( )  2   p 0 2 h A2 ( )  0 2

dA2 h ( )  0 dz 2 dA2 h ( )0 dz 2

(21)

4 Numerical examples 4.1 Analysis of simply supported FGM beam under sinusoidal pressure Let us consider a simply supported FGM beam subjected to a sinusoidal pressure as shown in Fig. 3. For convenience, the results are indicated in the coordinate system of Fig. 3 [6]. The Young’s modulus is assumed to be of the form E  E0 e αz h and the constitutive equation is the same as that of Ref. [7].

x

O

h l

z Figure 3:

A simply supported FGM beam subjected to sinusoidal pressure.

Two types of the material properties are considered, and the ratios of Young’s moduli of the top surface and the bottom surface are Eh E0 =10 and

Eh E0 =0.1. In the case of the former, the graded index  in Eqn. (5) corresponds to α  2.30 , and in the case of the latter, it corresponds to α=-2.30 . Further, in the case of the former, the load is applied on the softer WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)

126 Materials Characterisation V surface and in the case of the latter, on the harder surface. In both cases, Young’s modulus is E0 =1 [Gpa], Poisson’s ratio is ν  0.25 and the thickness is h  10 [mm]. The type of the applying force is sinusoidal pz ( x)  p0 sin ξx=p0 sin iπx l , ξ=iπ l , i=1,3,5 . Figure 4 shows the axial stress distribution at the middle span for Eh E0 =10 ,ξh=1 , and Fig. 5 shows the shear stress distribution at the middle span for Eh E0 =0.1 ,ξh=3 . In the figures, “Galerkin method” indicates the results of Ref. [6]. The vertical axis is the normalized stress values. The stress  xx is divided by the  xx ( x, h) , and the stress  xz is divided by the average value at the middle section. From these results, it can be noted that the present results agree well with the results by Ref. [6] for the various conditions of the analysis. Analysis of cantilever FGM beam under uniform pressure The analysis of the cantilever FGM beam subjected to the uniform pressure as shown in Fig. 2 have been carried out [9]. The length l is l = 1 [m] and the thickness h is h = 0.2 [m]. The material properties at z = z0 = 0 are given Table 1, and the magnitude of the transverse pressure p0 is p0 = 1 [Pa]. ξh=1, graded index=+2.30 Present solution

Galerkin method

σxx(x,z)/σxx(x,h)

2.0 0.0 0.0

0.2

0.4

0.6

0.8

1.0

-2.0 -4.0 -6.0 z/h

Figure 4:

Axial stress

 xx ( x, z ) /  xx ( x, h)

through the thickness of FGM

beam at x  l 2 . For the graded index component

 xx

  3 ,

Fig. 6 shows the distributions of the stress

at the clamped edge. The horizontal axis is the non-dimensional

coordinate z and the vertical axis is the stress values. In this figure, “Ref. [9]” means the results of Ref. [9]. The present results again are in good agreement with those of Ref. [9]. It was also confirmed that the distributions of the stress components  zz and  zx are also in good agreement. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


127

ξh=3 , graded index=-2.30 Present solution

Galerkin method

σxz/σxz(average)

2.0 1.5 1.0 0.5 0.0 -0.5 0

0.2

0.4

0.6

0.8

1

z/h

Figure 5:

Shear stress

 xz /  xz (average)

through the thickness of FGM

beam at x  l 2 . Table 1:

Materials properties.

Material constants

Values [1/Pa]

s

5.41  10 12

s

 1.51  10 12

0 11 0 13

0 s33

9.52  10 11

0 s44

1.37  10 10 graded index α=+3, -3

σxx(l, z)

Present (α=+3) Present (α=-3)

Ref.[9] (α=+3) Ref.[9] (α=-3)

250.0 200.0 150.0 100.0 50.0 0.0 -50.0 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0.1 -100.0 -150.0 -200.0 -250.0 z/h

Figure 6:

Stress at the clamped end

0.2

0.3

0.4

0.5

 xx (l , z ) (   3 ).


128 Materials Characterisation V 4.2 Analysis of simply supported FGM beam under uniform pressure The simply supported FGM beam as shown in Fig. 1 is analyzed. The sizes, the elastic constants and the applied load are the same as those of the previous section. The graded index α is assumed as α  3 . The distributions of the stress components σ xx at the x  l 2  0.5 and σ zx at the x  0 are plotted in Figs. 7– 8. The number of terms adopted in the Fourier series of Eqns. (9) and (14) is ten. We confirmed that taking at least eight terms is sufficient accuracy for the solution in this problem. graded index α=+3.0 20.0 10.0

σxx

0.0 -10.0 -0.5 -0.4 -0.3 -0.2 -0.1

0.0

0.1

0.2

0.3

0.4

0.5

-20.0 -30.0 -40.0 -50.0 -60.0 -70.0 z/h

Figure 7:

Stress σ xx through the thickness at x  0.5. graded index α=+3.0

5.0 4.5 4.0

σzx

3.5 3.0 2.5 2.0 1.5 1.0 0.5 0.0 -0.5

-0.3

-0.1

0.1

0.3

z/h

Figure 8:

Stress σ zx through the thickness at x  0 .


0.5


From these results, the value of the stress value stress  zx . Comparing the tensile stress compressive stress

 xx

 xx

 xx

129

is greater than that of the

on the bottom surface with the

on the top surface, the greater compressive stress is

caused on the softer surface. The shear stress distribution

 zx

is different from

that for homogeneous beams. The point where the maximum stress appears moves toward the softer surface. It may be possible to control distribution of stresses by adjustment of materials properties.

5 Concluding remarks In this paper, one analytical solution method for two-dimensional elastostatic problems in FGMs in which the properties of the materials vary exponentially in one direction is proposed. The method is based on the idea that the governing equation of this problem is of the same form as the governing equation for the plate bending with variable rigidity. The analysis method uses the Airy stress function. From analysis of a few numerical examples, most of the results correlate well with other solutions and the validity of the method is shown. The distributions of displacement components are not stated, however we can obtain them by integrating the strains and displacements relationships. Future work will be focused on development of an analytical method of solution for analyzing problems of FGMs having an arbitrary variation of their material properties expressed by functions other than exponential.

References [1] Suresh, S., and Moretensen, A., Fundamentals of Functionally Graded Materials, IOM Communications Ltd, London, 1998. [2] Chakraborty A., Gopalakrishnan S. and Reddy JN., A new beam finite element for the analysis of functionally graded materials, International Journal of Mechanical Science, Vol.45, pp.519-539, 2003. [3] Trung-Kien Nguyen, Karam Sab and Guy Bonnet, First-order shear deformation plate models for functionally graded materials, Composite Structures, 83, pp.25-36, 2008. [4] H.J. Xinag and J. Yang, Free and forced vibration of a laminated FGM Timoshenko beam of variable thickness under heat conduction, Composites: Part B Vol.39, pp.292-303, 2008. [5] Sankar BV., An elasticity solution for functionally graded beams, Composite Science Technology, 61, pp.689-696, 2001. [6] Zhu H. and Sankar BV., A combined Fourier series-Galerkin method for the analysis of functionally graded beams, Journal of Applied Mechanics, Vol.71, pp.421-424, 2004.


130 Materials Characterisation V [7] Sankar BV. and Tzeng JT., Thermal stresses in functionally graded beams, AIAA Journal, Vol.40, pp.1228-1232, 2002. [8] L.S. Miers and J.C.F. Telles, Two-dimensional elastostatic analysis of FGMs via BEFM, Engineering Analysis with Boundary Elements, Vol.32, pp.1006-1011, 2008. [9] Zheng Zhong and Tao Yu, Analytical solution of a cantilever functionally graded beam, Composite Science and Technology, Vol.67, pp.481-488, 2007. [10] Mansfield, E.H., The Bending and Stretching of Plates, Pergamon Press, pp.64-73, 1964.



131

Moment curvature analysis of concrete flexural members confined with CFRP grids A. Michael & P. Christou Department of Civil Engineering, Frederick University, Cyprus

Abstract The moment-curvature (M-Φ) diagrams define the maximum capacity of structural elements that are primarily subjected to bending moments and therefore their failure mode is flexural. The moment-curvature diagrams are also used to assess the ductility of structural elements and are therefore very important for the determination of the amount of plastic energy a structural element can absorb. One way of improving the reinforced concrete (RC) members is the confinement of concrete with FRP composites. The work presented in this paper includes the development of moment curvature diagrams for RC members using a fiber model. The RC members were designed as compression controlled members meaning that their failure initiates in the concrete prior to yielding of the steel tension reinforcement. These types of members have limited ductility. The introduction of a specific amount of Carbon Fibre Reinforced Polymer (CFRP) composite grid as confining reinforcement improves the ductility of the RC members by as much as 30%. The results from the fiber model are compared to available results from an experimental program conducted to evaluate the experimental improvement of the ductility of compression controlled members. The experimental and analytical results are a good match indicating that the fiber model is accurate and can be used to develop the moment-curvature diagrams of RC members confined with a CFRP composite grid. Keywords: CFRP grid, concrete confinement, moment curvature, ductility.

1 Introduction Compression controlled steel reinforced members, fail through crushing of concrete and therefore lack the ductility that is associated with yielding of the WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line) doi:10.2495/MC110121

132 Materials Characterisation V reinforcement. This is due to the fact that concrete as a material does not behave in a ductile manner. Concrete may exhibit limited ductility (also called pseudo ductility) but this ductility varies depending on the aggregate material and the strength of concrete used. High strength concrete is more brittle than low strength concrete and therefore for high strength concrete the pseudo ductility diminishes. This is similar to concrete members reinforced with Fibre Reinforced Polymer (FRP) composites. FRP reinforcing bars when loaded in tension, exhibit linear stress-strain behaviour up to rupture. There is no yield point and associated plateau to provide a ductile response when used as tensile reinforcement in concrete. Rooney and Taylor [1], Toutanji and Deng [2], and Grace et al. [3] found that the post cracking beam behaviour was linear to failure in concrete beams reinforced with glass FRP rods which shows the lack of a yield plateau and therefore the lack of a ductile response. There have been attempts to improve the ductility of FRP with little or no success. If these members are designed in such a way in order to fail through concrete crushing limited ductility may be exhibited much like in the case of compression controlled members. Therefore, the focus of improving member ductility should be on the concrete for compression controlled members. If the concrete in the compression zone of a flexure dominated member is confined then degradation of the compression zone at capacity is delayed, resulting in a more ductile response. Carbon Fibre Reinforced Polymer (CFRP) composite grids have been used to reinforce concrete decks and beams for both strength and crack control [4–8]. The CFRP composite grid has been used as confinement reinforcement with promising results. In a series of cylinder tests conducted by Michael et al. [9] it was found that the crushing strain of CFRP composite grid confined concrete was more than 2 times higher than the crushing strain of unconfined concrete. The advantage that the CFRP composite grid has over the use of FRP composite wraps that have been used for both confinement of concrete or strengthening of concrete members is the ability of the CFRP composite grid to be embedded in the concrete. The CFRP composite grid has openings that allow concrete to flow through thus can be embedded in the member during construction. This also provides environmental protection to the CFRP composite since the composite is not exposed to the natural elements that can cause environmental degradation to the FRP composite. Moment-curvature diagrams are used to assess the ductility of structural elements and are therefore very important for the determination of the amount of plastic energy a structural element can absorb. The research presented in this paper evaluated the moment-curvature response of compression controlled RC members confined with CFRP composite grid with the use of a fiber model that plots the moment-curvature diagram of such sections.

2 Experimental investigation The unique application of the CFRP composite grid as confining reinforcement was evaluated with the construction of two beams. Beam 1 had no CFRP composite grid tubes and served as the control beam and beam 2 had two CFRP WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


133

composite grid tubes placed in the compression zone. Details of the dimensions, amount and location of steel reinforcement and CFRP composite grid tubes are provided in Figures 1 and 2. The CFRP composite grid was formed into a circular tube using a 152 mm diameter plastic pipe and was held in place by a thin rope wrapped around the tube along its length. 5 in (127 mm)

3 in (76 mm)

2.75 in (70 mm) 18 in (457 mm) 6.65 ft (2.03 m)

2.7 ft (0.81 m)

6.65 ft (2.03 m)

16 ft (4.87 m) 1 in (25 mm)

1 in (25 mm)

#3 (#10) Stirrups

18 in (457 mm) 8 #10 (#32) Bars 2.75 in (70 mm)

2 in (51 mm)

3 in (76 mm) 12 in (305 mm)

Figure 1: 5 in (127 mm)

Details of control beam. 13 ft (3.96 m)

2.75 in (70 mm) 18 in (457 mm) 6.65 ft (2.03 m)

2.7 ft (0.81 m)

6.65 ft (2.03 m)

16 ft (4.87 m) 0.75 in (19 mm)

1 in (25 mm) 1 in (25 mm)

6 in (152 mm)

#3 (#10) Stirrups

18 in (457 mm) 8 #10 (#32) Bars 2.75 in (70 mm)

2 in (51 mm)

3 in (76 mm) 12 in (305 mm)

Figure 2:

Details of grid beam.


134 Materials Characterisation V The beams were tested in a simply supported four-point bending configuration (see Figure 3) in displacement control mode, that is, a constant displacement rate was applied independently of the amount of load. The span length between the two supports was 4.27 m, whereas the total length of the beams was 4.87 m. Actuator 1 Load Cell

Spreader Beam D15

D14

D13

D12

D11

D10

D9

D8

D7

D6

D5

D4

D3

D2

D1

18 in (457 mm)

1 ft (305 mm) Typ.

North Support

South Support 5.5 ft (1.7 m)

LVDT

3 ft (0.91 m)

Tilt Meter

Figure 3:

5.5 ft (1.7 m)

60 mm Foil Gauge

Test set-up.

3 Moment curvature analysis The moment-curvature (M-Φ) diagrams of the beams were calculated based on experimental data and also based on a theoretical fiber model. The experimental moment-curvature diagrams were calculated based on the displacement profile curves at increasing loads. The displacement profile equation was generated by fitting a 3rd order polynomial line to the displacement profile curves. The fitted lines were in good agreement with the displacement data (average R2 = 0.99). The polynomial displacement profile equation of the beam was used to determine the curvature at specific points along the length of the beam. The second derivative of the polynomial displacement profile equation is the curvature equation. The M-Φ curves of the beams were calculated using a fiber model. The compression zone of the tube beam was separated into two regions with regard to concrete properties. The first region was the concrete cover on top of the CFRP grid tubes and the second region was from the neutral axis to the top of the CFRP grid tubes (see Figure 4). The first region was extended approximately 0.25 in. (6.5 mm) below the top of the CFRP grid tubes to account for the unconfined concrete at the sides and between the grid tubes. The first region was assigned unconfined concrete properties and the second region confined properties. The concrete area from the neutral axis to the bottom of the grid tubes was unconfined but was treated as confined concrete for simplicity and because WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


135

the strains in that area close to the neutral axis were small and therefore the effect on the strength was minor. The unconfined region was divided into one rectangular layer while the confined region was divided into eight rectangular layers. For the control beam, the compression zone was treated as an unconfined concrete region and was divided into eight layers. 1 in. (25 mm)

c - 25 mm (c - 1 in.)

0.75 in. (19 mm)

Grid Tube N. A.

Unconfined concrete Confined concrete

Figure 4:

Tube beam cross-section fiber model.

Some of the assumptions employed in this fiber model include perfect bond between concrete and the reinforcing bars (strain compatibility), plain sections remain plain (linear strain distribution), the area below the neutral axis was considered cracked and was ignored in the force and moment calculation and monotonic loading and deformation of the section. The compressive force in each layer was calculated as the product of the area of the layer and the average stress. The average stress was determined based on the strain in the layer and the stress-strain curves of concrete. The Hognestad parabola was used to calculate the stress-strain curve of unconfined concrete. The modified Hognestad was used to calculate the stress-strain curve of confined concrete [10]. The parameters of the modified Hognestad were determined based on the available data from confined concrete cylinder tests [9]. The tensile force was calculated as the product of the area of steel reinforcement and the steel stress. Steel stress was determined based on the average strain of the two steel layers and by assuming elastic perfectly plastic stress-strain curve. The moment in the cross-section for the applied curvature distribution was determined by summing moments about the neutral axis. The moment curvature diagram was calculated by applying increasingly larger top concrete compressive strains (εc) to the cross-section and varying the location of the neutral axis (c) until equilibrium of the axial forces (zero) in the WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)

136 Materials Characterisation V cross-section for the applied strain was achieved. The curvature of the section (Φ = εc / c) was then calculated based on the concrete compressive strain and the location of the neutral axis. The experimental and fiber model M-Φ diagrams were calculated for both beams. The theoretical and fiber model M-Φ diagrams for the control beam (see Figure 5) were approximately the same. The fiber model M-Φ behaviour of the tube beam was comparable to the experimental behaviour (see Figure 6). The improvement in the ductility of the Curvature x 1000 (rads/m) 16 24 31

39

500

678

400

542

300

407

200

271 Experimental Fiber Model

100

136

0

0 0

0.2

0.4 0.6 0.8 Curvature x 1000 (rads/in)

Figure 5:

0

Moment (kip-ft)

Moment (kN-m)

8

8

1

M-Φ curves for control beam.

Curvature x 1000 (rads/m) 16 24 31

39

500

678

400

542

300

407

200

271

Expiremental Fiber Model

100

136

0

0

0

0.2 0.4 0.6 0.8 Curvature x 1000 (rads/in)

Figure 6:

M-Φ curves for tube beam.


1

Moment (kN-m)

Moment (kip-ft)

0


137

tube beam compared to the control beam, although relatively small, was evident when the experimental M-Φ curves for both beams were compared (see Figure 7). The area under the load-displacement curve represents the amount of energy a structure can absorb before failure. The amount of energy is also a good indicator of ductility. Ductile structures usually can absorb higher amounts of energy compared to non-ductile (brittle) structures.

8


39

500

678

400

542

300

407

200

271

Control Grid

100

Moment (kN-m)

Moment (kip-ft)

0

136

0

0

0

0.2 0.4 0.6 0.8 Curvature x 1000 (rads/in) Figure 7:

1

Experimental M-Φ curves.

The areas under the experimental moment-curvature curves, of both beams, were calculated and compared. The area under the load-displacement curve of the tube beam was approximately 37% more than that of the control beam. The energy up to peak load was elastic energy while the post peak energy was inelastic. Therefore, the 37% extra energy was primarily inelastic energy which indicated the ability of the tube beam to undergo inelastic deformation compared to the control beam. This represents an improvement in the ductility of the member and is due to the confining effect of the CFRP composite grid. Another way to measure ductility is through ductility factors. Curvature ductility factors are a good way to evaluate the ductility of RC members. The curvature ductility factor (μΦ) is defined as the ratio of the ultimate curvature (Φu) over the curvature at first yield (Φy). The definition of ultimate and yield curvatures is simple when the behaviour is elasto-plastic. Definition of these parameters is more complicated when the behaviour is not elasto-plastic. Usually the behaviour of reinforced concrete is not perfectly elasto-plastic. Therefore, the need for a consistent definition of the ultimate and yield curvatures was realized. Researchers have proposed definitions of ultimate and yield displacement and curvature as well as instructions on how to calculate them. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)

138 Materials Characterisation V The ultimate curvature has been defined as the curvature at which the lateral load in the descending portion of the load-displacement curve is not less than 80% of the maximum load (Pu) [11–14]. Park and Paulay [15] have argued that the available ultimate deformation (and therefore curvature) is not necessarily the deformation that corresponds to the maximum load capacity. They further argued that “when survival without collapse is the criterion, it is too conservative to define ultimate deformation as the deformation corresponding to the maximum load-carrying capacity. It would seem reasonable to recognize at least some of this deformation capacity after the maximum load has been reached and to define the available ultimate deformation as that deformation when the loadcarrying capacity has reduced by some arbitrary amount after maximum load. For example, a 10 or 20% reduction in maximum load-carrying capacity could be tolerated in many cases, but the exact amount would depend on the particular case.” The yield displacement (and therefore curvature) can be defined as the displacement at the intersection of the horizontal line representing the ideal lateral capacity, Pi, (nominal capacity using the ACI 318 approach and a reduction factor of unity) and the straight line that passes through zero and the point in the load-displacement curve at 75% of the ideal lateral capacity [11–13]. The definitions of yield and ultimate displacements are depicted in Figure 8. P Pu Pi 0.8 Pu 0.75 Pi

0

Figure 8:

d y or d c

y

du

u

d

Ductility factor component definitions.

The curvature ductility (μΦ) was approximately 1.5 and 2 for the control and tube beam respectively. The schematic for determining Φu and Φc for the control beam is shown in Figure 9. The same approach was used in determining the ductility factor for the tube beam. The curvature ductility factor of the tube beam was approximately 33% higher than the curvature ductility factor of the control beam. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


0

8


39 678

400

Mu Mi

542

300

0.8 Mu 0.75 Mi

407

200

271

100

136 c

0 0

Figure 9:

0.2

u

0.4 0.6 0.8 Curvature x 1000 (rads/in)

Moment (kN-m)

500

Moment (kip-ft)

139

0 1

Curvature ductility factor for control beam.

The curvature ductility factor using the fiber model moment-curvature diagram of the control beam was approximately 1.6 which was approximately 4% higher than the experimental. On the other hand, the curvature ductility factor of the tube beam was approximately 2.45 which was an overestimate of approximately 21%.

4 Conclusions In this paper the development of moment-curvature diagrams for compression controlled RC sections confined with CFRP composite grids was presented. These diagrams were compared with diagrams developed from experimental data for the same sections. Based on the results from this comparison the following conclusions can be drawn: 1. The results from the two beams indicate that the CFRP grid provided confinement to the concrete. Confinement was achieved using a series of small diameter CFRP grid tubes rather than wrapping the entire cross-section. 2. The curvature ductility factor of the beam utilizing the CFRP grid tubes was improved by 33% compared to the control beam with no grid tubes. The improvement was significant, considering the small amount of CFRP that was used to confine the compression zone. 3. The energy dissipation of the tube beam was 37% higher compared to the controlled beam due to the effect of concrete confinement. This increase in the energy dissipation of the element also represents an improvement in the ductility due to concrete confinement from the CFRP grid tubes. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)

140 Materials Characterisation V 4.

5.

The results from the fiber model are fairly accurate compared to the experimental data and therefore the model is considered valid. However, there is a divergence after the peak load and further refinement of the model is required. The number of test specimens is limited and therefore the more testing is required. Due to the small number of test specimens the conclusions are drawn with great caution.

References [1] Rooney, J., and Taylor, S., 2004, “Flexural Behaviour of Steel and GFRP Reinforced Concrete Beams,” Composites: Part B, Vol. 27, No 3-4, pp. 38-40. [2] Toutanji, H., and Deng Y., 2003, “Deflection and Crack Width Prediction of Concrete Beams Reinforced with Glass FRP Rods,” Construction and Building Materials, Vol. 17, No 1, pp. 69-74. [3] Grace, N. F., Soliman, A. K., Sayed, G. A., and Saleh, K. R., 1998, “Behavior and Ductility of Simple and Continuous FRP Reinforced Beams,” Journal of Composites for Construction, Vol. 2, No 4, pp. 186-194. [4] Rahman, A. H., Kingsley, C. Y., and Kobayashi, K., 2000, “Service and Ultimate Load Behavior of Bridge Deck Reinforced with Carbon FRP Grid,” Journal of Composites for Construction, Vol. 4, No 1, pp. 16-23. [5] Yost, J. R., Goodspeed, C. H., and Schmeckpeper, E. R., 2001, “Flexural Performance of Concrete Beams Reinforced with FRP Grids,” Journal of Composites for Construction, Vol. 5, No 1, pp. 18-25. [6] Harries, K. A., and Gassman, S. L., 2003, “Load Tests of Reinforced Concrete Catch Basing Knockout Panels,” Department of Civil and Environmental Engineering, University of South Carolina, Report No ST03-01, p. 21. [7] Shao, Y., Johnson, C., and Mirmiran, A., 2003, “Control of Plastic Shrinkage Cracking of Concrete with TechFab Carbon FRP Grids,” Department of Civil, Construction, and Environmental Engineering, North Carolina State University, Report to Tech-Fab Inc, p. [8] Hamilton, H. R, Cook, R. A., and Alfonzo, L., 2006, “Crack Control in Toppings for Precast Flat Slab Bridge Deck Construction,” Department of Civil and Coastal Engineering, University of Florida, Engineering and Industrial Station, Final Report Submitted to FDOT, UF Project No 00030907, p. 120. [9] Michael, A. P., Hamilton, H. R. III, and Ansley, M. H. (2005). “Concrete confinement using carbon fiber reinforced polymer grid”, American Concrete Institute, SP230-56, Vol. 230, pp. 991-1010. [10] Park, R., and Paulay, T., 1975 (a), “Chapter 6: Ultimate Deformation and Ductility of Members with Flexure,” Reinforced Concrete Structures, John Wiley & Sons, New York, NY, pp. 195-269.



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[11] Priestley, M. J. N., and Park, R., 1987, “Strength and Ductility of Concrete Bridge Columns under Seismic Loading,” ACI Structural Journal, Vol. 84, No 1, pp. 61-76. [12] Zahn, F. A., Park, R., and Priestley, M. J. N., 1990, “Flexural Strength and Ductility of Circular Hollow Reinforced Concrete Columns without Confinement on Inside Face,” ACI Structural Journal, Vol. 87, No 2, pp. 156-166. [13] Sheikh, S. A., and Khoury, S. S., 1993, “Confined Concrete Columns with Stubs,” ACI Structural Journal, Vol. 90, No 4, pp. 414-431. [14] Yeh, Y. K., Mo, Y. L., and Yang, C. Y., 2002, “Seismic Performance of Rectangular Hollow Bridge Columns,” Journal of Structural Engineering, Vol. 128, No 1, pp. 60-68. [15] Park, R., and Paulay, T., 1975 (b), “Chapter 11: Strength and Ductility of Frames,” Reinforced Concrete Structures, John Wiley & Sons, New York, NY, pp. 496-609.




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Application of effective media theory in the characterization of the hygrothermal performance of masonry Z. Pavlík, E. Vejmelková, L. Fiala, M. Pavlíková & R. Černý Department of Materials Engineering and Chemistry, Faculty of Civil Engineering, Czech Technical University in Prague, Czech Republic

Abstract Hygric and thermal properties of several types of materials which are used in reconstructions of historical stone masonry on Czech territory are experimentally measured at first. The obtained data are then subjected to a homogenization procedure, and the effective heat and moisture transport and storage parameters of the typical fragment of masonry are calculated. For the homogenization, a mixing model originally derived for applications in dielectric studies was employed taking into account the theoretical bounds of the investigated material parameters. The obtained data can find use in damage assessment of historical masonry using methods of computational simulations and analysis. Keywords: hygric properties, thermal properties, historical stone masonry, argillite, sandstone, lime-based plaster, homogenization techniques.

1 Introduction The hygrothermal behaviour of climatically exposed components and structures of historical buildings is related to the hygric, thermal, mechanical and other physical and chemical properties of inbuilt materials. Their knowledge represents necessary information for the understanding of the building performance as well as the first step in avoiding damage, degradation or the undue heat loss from the constructions. The damage assessment of historical masonry and other structures due to the negative effects of moisture and temperature can be done effectively by means of mathematical and computational modelling. In this way, the time development of WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line) doi:10.2495/MC110131

144 Materials Characterisation V moisture and temperature fields can be obtained which is a crucial for a proper assessment of possible future damage [1]. Their prediction is a very important task when preserving historical bridges and buildings or insulating existing buildings and components. The moisture and temperature values can be then assigned to the mechanical properties and to the risk of consequent damage. For instance Charles Bridge in Prague, Czech Republic, which was subjected to reconstruction works lately, is a typical example of extensive damage mostly brought about by temperature and moisture impacts. For the effective application of computational modelling in the buildings’ performance analysis, there is necessary complete knowledge of material properties of applied materials as well as of initial and boundary conditions of simulations. As the initial conditions for the hygric and thermal analysis, the moisture and temperature fields experimentally measured in the investigated structure can be used. Boundary conditions of the computer simulations are given by the climatic loading that can be simply adopted using meteorological data. In the technical practice, simplified calculations of building structures’ behaviour are very often used. Here, in the computational simulations of heat and moisture transport processes, brick or stone masonry is often understood as a single material, the brick or stone itself. This simplification may work reasonably well when the properties of the brick or stone and the mortar are similar. However, this may not always be the case in some historical buildings, where low-quality mortars were often used [2]. On that account, the results of such calculations are diverse from the real state of the buildings, and cannot be effectively used for estimation of some appropriate materials’ damage and problems in hygrothermal behaviour. Therefore, solution of this problem must be found in order to improve the accuracy and reliability of computational modelling of building structures performance. There are two ways how to meet the above given requirements on reliability and accuracy of computational simulation of heat and moisture transport. The first possibility is the use of sophisticated simulation tools that allow discretization of studied structure in such resolution, that all the properties of inbuilt materials can be assigned to the specific places of the structure. Although this procedure is generally correct, it is too complex and sophisticated for wider application in engineering practice, especially in case of conservationists of historical buildings that are usually well oriented in materials’ problems without knowledge of methods of computational modelling. The second, and probably the most straightforward way to solve the problem of differences in the properties of the particular components of masonry is the utilization of the effective media theory. This procedure we introduce in the presented work as an effective tool for simplification of computational modelling of transport processes in the masonry. Application of homogenization principles produces the macroscopic equations which may be used when analyzing the masonry as a whole. In these equations, the effective parameters are used instead of the parameters of the brick or stone.



145

Quite apparently, the homogenization process cannot be done without the exact knowledge of the properties of all materials constituting the masonry and of the amount of each material in the analyzed wall. Therefore, the experimental measurement of all the material parameters characterising the process of heat and moisture transport and accumulation in inbuilt masonry materials is necessary at first. These data are than implemented in the mixing models based on effective media theory, and the effective thermal and hygric parameters of the whole masonry can be calculated.

2 Studied materials Two types of stones that were in the past often used in Czech region as bearing masonry materials are analysed together with lime-based mortar with pozzolana admixture. Many historical buildings in the Czech Republic were built using similar kinds of sandstone. Siliceous raw-grained sandstone was usually used for historical architectural constructions (walls, portals, window frames) for its strength. Ornamental parts of the architecture (gothic flowers, romantic shells) and sculptures (from the Romanesque period up to now) were made of finegrained calcite-argillaceous sandstone. In this work, sandstone from Mšené-lázně quarry, Czech Republic, is chosen. It is fine-grained psamitic equigranular rock, about 95% of which is made up of suboval quartz clasts. Other mineral grains are present only as minorities (tourmaline, epidote, muscovite and zircone). Quartz grains reach up to 0.1 mm in diameter, but those of muscovite are larger, up to 0.3 mm [3]. The matrix is formed by clay minerals (mainly kaolinite). Also the argillite was very popular material in historical architecture. It was used for sacral as well as for secural buildings, flagstone pavements, roof slabs, and facing. The studied argillite is coming from quarry Džbán, Czech Republic. Its main constituents are illite, calcite, minerals on the basis of SiO2 having granularity 0.3 – 0.15 mm, feldspar, and mica, whereas rigid materials form 40 – 60% of argillite volume. Within the reconstruction of historical buildings, renewal and restoration of interior and exterior plasters50and mmortars is often done. From the point of view of a historian, it is not acceptable to use lime-cement mortars in Romanesque, Gothic, Renaissance, and Baroque buildings. However, the pure lime-based plaster does not exhibit sufficient resistivity against moisture action. On that account, proper hydraulic admixtures must be used that enhance the durability of mortars. These materials must have similar composition as the historical materials and they have to be applicable by the original technological processes. As the chemical analyses of many plasters from historical buildings show, the past centuries external plasters that are preserved until today contain products formed by lime reaction with pozzolanic or hydraulic admixtures. Pozzolanic admixtures appeared to have positive effect on properties of lime binder in the past. On that account, lime-based mortar with metakaolin addition as pozzolana admixture was chosen. The composition of the lime-metakaolin plaster was as


146 Materials Characterisation V follows: hydrated lime – 400 g, metakaolin - 80 g, natural quartz sand with continuous granulometry 0 to 2 mm – 1 440 g, and water – 480 g.

3 Experimental methods Basic material properties of all tested materials were determined at first. Bulk density and matrix density were measured using gravimetric method and helium pycnometry, and then total open porosity was calculated. The samples’ dimensions for these measurements were 40 x 40 x 20 mm. For determination of water vapour transmission properties we applied cup method, according to the European standard EN ISO 12571 [5]. The measurements were carried out in steady state under isothermal conditions. We used dry cup arrangement of the experiment. Here, the sealed cup with the studied material sample containing burnt CaCl2 (0% relative humidity) was placed in a controlled climatic chamber at 25±0.5°C and 50% relative humidity and it was weighed periodically. The circular samples had diameter 95 mm and thickness of 20 mm. The steady state values of mass gain were utilized for the determination of the water vapour permeability. On the basis of measured water vapour permeability, the water vapour diffusion coefficient and water vapour resistance factor were calculated using the simple formulas given in [6, 7]. For determination of moisture diffusivity as function of moisture, the moisture profiles were measured at first. The measurements were done on samples having dimensions of 40 x 20 x 300 mm in 1-D experimental arrangement of water transport. The moisture content in specific position in sample was measured by capacitance technique calibrated by gravimetric method. The moisture dependent moisture diffusivity was then calculated using inverse analysis of measured moisture profiles by means of Boltzmann-Matano treatment [8]. In the sorption isotherm measurement, the samples were placed into the desiccators with different salt solutions to simulate different values of relative humidity. The mass of samples was measured in specified periods of time until steady state value of mass was achieved. Then, the volumetric moisture content was calculated and sorption isotherm of each tested material was plotted. The thermal conductivity as the main parameter of heat transport was determined using the commercial device ISOMET 2104 (Applied Precision, Ltd.). ISOMET 2104 is a multifunctional instrument for measuring thermophysical parameters which is based on the application of an impulse technique and is equipped with various types of optional probes. Thermal conductivity was measured in the moisture range from the dry state to full water saturation on the 70 mm cubes.

4 Effective media theory and homogenization techniques Application of effective media theory allows determination of the properties of the whole masonry or building structure instead of parameters of the particular inbuilt materials. In terms of effective media theory, the final composite structure WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


147

(in our case stone masonry wall) can be considered basically as a mixture of walling blocks and mortar. In more precise calculations, each of these materials can be further considered as a mixture of three phases, namely solid, liquid and gaseous phase (in four phase systems, the effect of bound water can be included) that form their matrix and porous space. There are two basic approaches that can be applied for determination of thermal and hygric properties of the stone masonry. The first possibility is to apply homogenization techniques on the moisture dependent material data of the materials involved in the masonry. This simplified procedure was used in this work. The second possibility is based on the complete knowledge of material properties of the particular components forming the porous body of the structure (dry stone, dry mortar, free water, bound water and air). From the properties of particular components and their volumetric fractions the effective properties of the stone masonry can be accessed. On the basis of previous experience with application of homogenization techniques, we have chosen for the calculations performed in the presented work original Lichtenecker’s formula that was proved to be satisfactory for evaluation of moisture dependent thermal and hygric properties of porous building materials [9]. Also the application of four phase models looks promising. However, in case of the masonry studied in this work, determination of amount of bound water represents very complex problem. The Lichtenecker’s equation adjusted for the studied masonry wall assumes that the effective hygric and thermal parameter of the considered material satisfies equation

peff  f b pb  f m pm , k

k

k

(1)

where peff is calculated effective parameter of masonry, fb volumetric fraction of walling blocks in masonry, fm volumetric fraction of mortar, pb measured material parameter of walling block, pm measured parameter of mortar, and k is free parameter describing basically the path from the anisotropy at k = -1.0 to another anisotropy at k = 1.0. However, the Lichtenecker’s equation may be also applied for isotropic materials. On the basis of previous experience, we have used the value of the parameter k = 0. The effective parameters of multi-phase material cannot exceed the bounds given by the parameters of particular fractions of its constituents. Here, the Wiener bounds according to the Wiener’s original work were used [9]. These bounds can be expressed by the following relations

peff 

1 n

 i 1

fi pi

,


(2)

148 Materials Characterisation V n

peff   f i pi ,

(3)

i 1

where Eq. (2) represents the lower limit and (3) the upper limit of the investigated effective material parameter (fi is the volumetric fraction of the particular phase, in our case argillite, sandstone and mortar), pi its material parameter).

5 Studied masonry For application of homogenization theory for the evaluation of the effective hygric and thermal parameters of masonry, the typical fragment of the wall was constructed in a simplified way similar to common brick masonry as shown in Figure 1.

Figure 1:

Scheme of the reference masonry wall (dimensions in mm).

Using this scheme, the volumetric fractions of the particular materials in the wall were calculated. The volumetric fraction of walling materials is equal to 0.824, whereas the volumetric participation of mortar is 0.176. Within the calculations, effective parameters of two different walls were accessed. The studied walls consisted always of lime based metakaolin mortar, and of one walling material, nominally sandstone or argillite.

6 Results and discussion Basic properties of all tested materials are summarized in Table 1. Each result represents the average of five measured values. All the studied materials have WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


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proved high porosity, what is very positive factor from the point of view of their presumed application in reconstruction of historical buildings. Water vapour transmission properties of masonry materials are presented in Table 2. We can see systematically very low values of water vapour resistance factor what is again very prospective finding for application of all tested materials in historical masonry. Since the historical masonry usually exhibits increased moisture content, there is necessary to apply within the reconstruction processes such materials that will allow moisture evaporation from the renovated structures within the warm periods of the year. Table 1: material sandstone argillite lime-metakaolin mortar Table 2:

Basic material properties of masonry materials. bulk density (kg/m3) 1807 1353 1690

matrix density (kg/m3) 2627 2235 2620

total open porosity (-) 0.31 0.39 0.35

Water vapour transmission properties of masonry materials.

material

sandstone argillite lime-metakaolin mortar

dry cup arrangement, 0 – 50% water vapour water vapour water vapour permeability (s) diffusion resistance factor coefficient (m2/s) (-) 2.4E-11 3.3E-06 7.0 2.9E-11 4.1E-06 5.7 1.8E-11 2.5E-6 10.1

In Table 3, there are presented the effective diffusion parameters of studied stone masonry calculated on homogenization principles. For verification of calculated results, also the limiting bounds of effective parameters are presented. Looking at the obtained data of water vapour transmission properties, one can see that all the effective parameters lie between the Wiener’s bounds which basically verifies their correctness. Effective moisture diffusivity as function of moisture content is given in Figures 2, 3. We can see high dependence of this moisture transport parameter on moisture content. This materials’ behaviour significantly affects the hygrothermal performance and consequently the durability. Systematically, the highest moisture diffusivity exhibits sandstone. Its highly porous structure formed by high radius pores allows fast liquid moisture transport in comparison with argillite that is characteristic also by high total open porosity, but its pore size is much smaller and structure more fine-grained.


150 Materials Characterisation V Table 3:

Water vapour transmission properties of stone masonry.

Sandstone wall – homogenization, dry cup arrangement – 0-50% RH water vapour permeability (s)

water vapour water vapour diffusion resistance coefficient (m2/s) factor (-)

Wiener’s lower bound

2.29E-11

3.11E-06

7.40

Wiener’s upper bound

2.24E-11

3.05E-06

7.55

Lichtenecker model, k=0

2.27E-11 3.08E-06 7.47 Argillite wall – homogenization, wet cup arrangement – 0-50% RH water vapour permeability (s)

water vapour water vapour diffusion resistance coefficient (m2/s) factor (-)

Wiener’s lower bound

2.74E-11

3.73E-06

6.17

Wiener’s upper bound

2.61E-11

3.55E-06

6.47

Lichtenecker model, k=0

2.69E-11

3.65E-06

6.30

Figure 2:

Effective moisture diffusivity of sandstone masonry.

The calculated effective sorption isotherms of studied masonry are given in Figures 4 and 5. Typically, the highest capacity for water vapour adsorption can be observed for lime-metakaolin mortar. However, in hygroscopic moisture range, the highest values of adsorbed moisture were measured for argillite. The data obtained for sandstone were systematically the lowest. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


Figure 3:

Effective moisture diffusivity of argillite masonry.

Figure 4:

Effective sorption isotherm of sandstone masonry.


151


Figure 5:

Effective sorption isotherm of argillite masonry.

Measured and calculated thermal conductivity of all studied materials and two types of masonry is presented in Figures 6 and 7. The values of effective thermal conductivity in dependence on moisture content are systematically between the range of Wiener’s bounds, which basically proves reliability of applied homogenization technique.

Figure 6:

Effective thermal conductivity of sandstone wall.



Figure 7:

153

Effective thermal conductivity of argillite wall.

7 Conclusions An example of application of homogenization technique for determination of hygric and thermal parameters of the stone masonry was introduced in the paper. The obtained results indicate that the Lichtenecker’s equation may be successfully used for such type of applications. Nevertheless, there is still an open task to verify the reliability of applied homogenization model by laboratory experiments that should confirm or controvert the presented results.

Acknowledgement This research has been supported by the Czech Ministry of Education, Youth and Sports, under project No MSM 6840770031.

References [1] Pavlíková, M., Pavlík, Z. & Černý, R., Hygric and Thermal Properties of Materials Used in Historical Masonry. Proc. of the 8th Symposium on Building Physics in the Nordic Countries, Technical University of Denmark: Lyngby, pp. 903-910, 2008. [2] Vejmelková, E., Pavlík, Z., Fiala, L., Pavlíková, M. & Černý, R., Heat and Moisture Transport Properties of Stone Masonry Materials. Proc. of the Building Physics Symposium, Katolieke Universiteit Leuven, Leuven, pp. 113-116, 2008.


154 Materials Characterisation V [3] Pavlík, Z., Michálek, P., Pavlíková, M., Kopecká, I., Maxová, I. & Černý, R., Water and Salt Transport and Storage Properties of Mšené Sandstone. Construction and Building Materials, 22(22), pp. 1736-1748, 2008. [4] Pavlík, Z., Identification of parameters describing the coupled moisture and salt transport in porous building materials, CTU Prague, p. 145, 2009. [5] EN ISO 12572, Hygrothermal performance of building materials and products, determination of water vapour transmission properties, the European Committee for Standardization, Brussels, 2001. [6] JIřičková M., Application of TDR Microprobes, Minitensiometry and Minihygrometry to the Determination of Moisture Transport and Moisture Storage Parameters of Building Materials, CTU Prague, p. 102, 2004. [7] Roels S., Carmeliet, J., Hens, H., Adan, O., Brocken, H., Černý R., Pavlík Z., Hall Ch., Kumaran K., Pel L. & Plagge R., Interlaboratory Comparison of Hygric Properties of Porous Building Materials. Journal of Thermal Envelope & Building Science, 27(4), pp. 307-325, 2004. [8] Roels, S., Carmeliet, J., Hens, H., Adan, O., Brocken, H., Černý, R., Pavlík, Z., Hall, C., Kumaran, K., Pel, L., A Comparison of Different Techniques to Quantify Moisture Content Profiles in Porous Building Materials. Journal of Thermal Envelope & Building Science, 27(4), pp. 261-276, 2004. [9] Pavlík, Z., Vejmelková, E., Fiala, L., Černý, R., Effect of Moisture on Thermal Conductivity of Lime-Based Composites. International Journal of Thermophysics, 30(6), pp. 1999-2014, 2009.



155

3D FIB reconstruction and characterisation of a SOFC electrode S. Chupin, N. Vivet, D. Rochais & E. Bruneton CEA-Le Ripault, France

Abstract SOFC (Solid Oxide Fuel Cells) appear to be a great alternative way to produce electricity from hydrogen with high efficiency and no greenhouse gas emissions. SOFC are efficient at high temperatures (around 800°C) and are meant to be used for stationary applications as heat and electric co-generation devices. To understand how gases, electricity and heat flow through these media and to improve their efficiency, it is critical to know the actual microstructure of these electrodes. The studied electrodes are Ni-YSZ cermets in which characteristic element sizes are around 1m. The 3D microstructure has been reconstructed using FIB (Focused Ion Beam) tomography. This technique has been used on several samples (different Ni-YSZ proportions) and gives representative 3D volumes of around 10×10×10µm with a 10nm resolution. Theses 3D volumes are then analysed to extract some important structural parameters such as volumetric proportions, active surfaces, connectivity of each components and “three phase boundaries” (TPB). Then, the 3D reconstructed volumes have been used to determine homogeneous media equivalent properties such as thermal, ionic and electric conductivities. These homogeneous equivalent properties are estimated using a hot guarded plate simulation that takes into account each component properties and the 3D structure. Keywords: SOFC, FIB, tomography, thermal conductivity.

1 Introduction Solid oxide fuel cell (SOFC) based-technology is one of the most promising energy conversion systems due to its high efficiency and fuel flexibility [1]. The materials involved in their constitution have become increasingly sophisticated, both in composition and microstructure. One of the most common anode WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line) doi:10.2495/MC110141

156 Materials Characterisation V materials for SOFC is a porous Ni-YSZ cermet. The most important requirements for the anode are catalytic activity, stability, ionic conductivity, compatibility and porosity. For most of these conditions some experimental studies can be investigated to estimate the corresponding parameters. Recently, 3D reconstructions of SOFC electrodes have been performed by X-ray computed tomography (XCT) [2, 3] and mainly by focused ion beam - scanning electron microscopy (FIB-SEM) [4–7]. FIB tomography principle consists in ablating a structure physically by FIB slicing (Ga+ ions) followed by a digital reconstruction since SEM images are collected after each ablation step. The obtained images are then aligned and superimposed in order to generate the volume [8]. In the present study, the optimization approach to obtain a high quality 3D reconstruction of a Ni-YSZ anode using FIB-SEM is described first. Then, from these 3D data, various microstructural parameters, and interfacial parameters are quantified. Moreover, by solving the diffusive transport equation on the analysed volume the effective thermal, electrical and ionic conductivities of the sample have been estimated.

2

Experimental

2.1 Preparation of the anodic cermet The deposition of the analysed NiO-8YSZ anode has been performed by screenprinting onto circular 8mol.% yttria doped-zirconia (8YSZ) supports using ink with optimized composition and viscosity. The anode is a three layers coating on the electrolyte substrate. NiO from J.T. Baker (USA) and 8YSZ from Tosoh (Japan) have been used as raw materials to prepare the NiO-8YSZ cermet at different Ni volume ratio. The preparation has been done to expect the 8YSZ/Ni volume ratio to be 60/40.

Figure 1:

Schematic representation of the experimental method used to obtain the 3D reconstruction of the SOFC anode from FIB tomography.



157

2.2 FIB procedure The image set has been collected with a FEI dual-beam Strata 400-S FIB system. The procedure to obtain the 3D data from the initial sample is schematically presented in Figure 1. 2.2.1 Preparation of the sample and of the region Of interest (ROI) In order to easily distinguish the pores during SEM observation and to avoid mistakes due to the depth of field in the SEM images, the sample has been infiltrated with an epoxy resin under vacuum. To protect the sample from accidental ion milling and erosion, a representative cross sectional area of the sample has been located by electron imaging and coated with a Pt protective layer of 2-3µm (Figure 2) using an in-situ metallorganic ion source. Above all, Pt deposition associated with sample infiltration and surface polishing were necessary to avoid, or at least strongly reduce, a common artefact so-called “waterfalling” or “curtaining” effect. The FIB has been used to mill wide and deep trenches around the ROI with a maximal aperture current of 21nA, constituting a “U-shaped trench” [9]. A quite high dimension free space around the ROI is essential to prevent re-deposition of the sputtered material during the sectional milling process. If the space around the ROI is too small, images with high concentration of artefact features or uneven brightness are obtained.

Figure 2:

SEM image of the sample before the “milling and imaging” procedure. The stage tilt is kept constant at 52° from the electron beam while the sample surface (xz plane) is perpendicular to the ion beam (y direction). The region of interest (ROI) is displayed in the orange rectangle. The sample is sliced in the z direction. The small rectangle on the Pt deposition (reference mark) is used to measure the interslice (z) and correct from drifting effects.


158 Materials Characterisation V 2.2.2 Milling and imaging process Since the “milling and imaging” procedure is very time consuming, uncorrected drift of electron beam, stage and sample can greatly affect the interslice real value. For this reason a correction procedure has been used [10]. The SEM and FIB operating conditions for the sequential ion milling and electron imaging have been optimized. The FIB milling has been performed with an ion-beam current of 460 pA at 30 kV. For SEM imaging, a magnification of 12 kX and a Through the Lens Detector (TLD) operating in BSE mode with low scan rate have been used. As the difference between the backscatter coefficients of Ni and 8YSZ are maximum for accelerating voltage smaller than 1kV, a 0.5kV accelerating voltage has given an optimal contrast between Ni (bright), 8YSZ (grey) and pores (dark) (Figure 3(a)). In this case, 115 images have been recorded using a manual procedure along the sample thickness (Figure 3(a)). 2.2.3 Data processing Data processing includes the following steps: i) alignment of the consecutives slices, ii) correction of the dimensions taking into account the tilt of the electron beam, iii) thresholding of the grey levels and labelling of phases, iv) resampling of the data to obtain cubic voxels and v) 3D image generation. Most of the processing steps have been performed using the AVIZO 6.2 software [11].

Figure 3:

Images of:(a) one of the original 115 serial-sections collected with a BSE detector, so that Ni appears brighter than 8YSZ and that the impregnated pores appear dark, (b) final 3D image after alignment, delineate, cropping and adjustment of voxel size operations showing Ni (bright), 8YSZ (grey) and pore (dark) phases.

The absolute dimensions in the x direction are obtained from a calibrated SEM magnification. Because of oblique SEM imaging at an angle of 52°, distances in the y direction have been corrected. The 3D image dimensions is finally 8.66×9.79×11.41µm3 (volume= 967.36µm3 corresponding to about 1 billion voxels).



159

3 Quantification results and discussion First, volume fractions of each phase and their corresponding connectivity across the analysed volume have been determined. The interfacial planes surface area between two phases and the three phase boundaries lines length have been evaluated. Then, effective thermal, electronic and ionic conductivities have been estimated by solving numerically the diffusive transport equation. The most part of this quantitative analysis has been performed with home made C programs on a personal computer [dual CPU Intel® Xeon® E5440 @ 2.83GHz, 32 GB RAM]. 3.1 Volume fraction The volume fractions calculated from 3D data have given the following results: 41% for pores, 33% for 8YSZ and 26% for Ni. The solid volume fractions have been estimated to 44% Ni and 56% 8YSZ in the studied anode, which are in good agreement with the expected values (40 vol.% Ni / 60 vol.% 8YSZ). This result provides evidence of the good representativity of the analysed volume. 3.2 Volumetric connectivity In order to check the connectivity of each phase, a home made program based on a Hoshen-Kopelman algorithm has been used [12]. The voxels linked together by a face form a cluster [13]. When two voxels are only linked by a vertex or an edge, they are not considered to be connected. Finally, if a cluster is connected to the six boundary faces of the studied volume it is defined as “percolated” (Figure 4).

Figure 4:

Ni clusters representation showing that the main part (87.4%) of the Ni phase is percolated, i.e. in contact with the six boundary faces of the volume (in yellow).


160 Materials Characterisation V This analysis reveals that the three phases are mainly constituted of a very large “percolated” cluster which represents 99.8%, 99.1% and 87.4% of pore, 8YSZ and Ni total volume respectively. This high percolation of all the three phases suggests good transport ability for gases in the pores, as well as for charge carrier in the solid phases. 3.3 Interface properties Electrochemical and catalytic activities of an electrode are generally described by only one parameter: the TPBL which is correlated to the number of regions where the electrochemical reactions is able to take place. However not all of them may be active due to non-contiguous regions (isolated cluster). So, the specific surface and interface areas have been calculated only inside the percolated volumes. In the pie diagram of Figure 5, the specific interface areas are indicated. Thus, about 50% of the total nickel surface is exposed to the porous phase and can be used for surface catalytic reactions with the fuel.

Figure 5:

Pie diagram representing the specific surface areas of pores, 8YSZ and Ni and the specific interface areas between the three neighbouring phases (surface and interface are normalized by volume sample). The active TPBL is 7.4µm.µm-3.

The TPB are the regions where the three phases (pores, 8YSZ, Ni) meet each other. The TPB analysis procedure has been apply first to the initial volume and yields to a TPB density of 11.2µm.µm-3. The “activity” of a TPB is subject to numerous definitions in the literature. The major part of the authors defined a TPB as “active”, if it lies three percolated phases (pore, Ni, 8YSZ) [4, 7]. The “active” TPBL has been calculated and equal to 7.4µm.µm-3. This value is relatively high compared to those reported in the literature. This suggests a high electrochemical performance for the studied electrode. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


161

3.4 Effective conductivities calculations 3.4.1 Effective diffusion coefficient calculation In the absence of external sources and in the case of the steady state condition, the distribution of  (for all conduction or diffusion phenomena) in the material is governed by the diffusive transport equation [14]:

  .  D   0





(1) D is defined as the diffusion coefficient in the material. Each voxel is considered uniform with a given diffusion coefficient depending of its phase (Dvoxel. Table 1). This equation is solved using an implicit finite difference method based on a conjugate gradient algorithm [14]. Different Dirichlet boundary conditions are fixed on two opposite faces (    1 on one face and

  2

on the opposite). Null Neumann boundary conditions are imposed on

the other faces. Once the  distribution within the volume is obtained (Figure 6), the average flux (Φ) over a cross section of surface area S can be calculated:

  Φ    D voxel  θ . n dS

(2)

S Table 1:

Thermal, ionic and electronic conductivities at 1123 K of pores, 8YSZ and Ni data [17–21].  (W.m .K )  ionic (10-3 S.cm-1)  electronic (S.cm-1) -1

Figure 6:

-1

Pores (H2) 0.48 0 0

8YSZ 2 14 10-6

Ni 50 10-5 104

Reconstructed 8YSZ-Ni SOFC Anode and the electric field resulting of a 1V potential difference between two opposite faces.


162 Materials Characterisation V For the equivalent homogeneous material, the average conduction flux Φ can be written [14]:

Φ  Deff S

θ1  θ 2 

(3)

xi max

The combination of eqns (2) and (3) leads to estimate Deff as:

Deff 

 θ    Dvoxel  xi S θ1  θ 2  S  xi max

  dS 

(4)

At 1123K and under one atmosphere the mean free path of H2 is around 900nm, and is bigger than our voxel size [15]. Then the concentration cannot be considered uniform in a voxel and the bulk diffusion assumption is not valid anymore. In the case of electron, ion and heat transport the corresponding mean free path are smaller than 100nm. So, only heat, ion and electron diffusive transport have been calculated in this work. The results are given in Table 2 where the different values of Deff are reported under the form of thermal, electronic and ionic effective conductivities. The effective thermal conductivity appears quite isotropic with a slighter lower value in the z direction, close to 4 W.m-1.K-1. This value has the same order than those generally reported for SOFC anodes [16]. Table 2:

Thermal, electronic, ionic effective conductivities in the x, y and z directions from the calculation of Deff at T= 1123 K.  (W.m .K )  ionic (10-3 S.cm-1)  electronic (S.cm-1) -1

-1

x 4.23 1.43 281

y 4.54 1.19 252

z 3.27 1.03 66

The effective electronic conductivity appears anisotropic: the electronic value in the z direction is smaller than the ones in the x and y direction. The evolution of the Ni concentration along each axes has shown that the Ni proportion along z in more important than along the other axes. It creates a stricture effect that decreases the effective conductivity. Experimental four-point electrical measurements performed at 1123 K on the anode surface lead to a value of the same range: 81 S.cm-1. Ionic conductivity in the studied anode is quite isotropic and close to 10-3. As a value of 1.4 10-2 had been used for the calculation, the microstructural effect for the loss of conductivity is closed to a factor 10.

4 Conclusion This work has focused on the 3D analysis of the microstructure in a Ni-8YSZ SOFC anode reconstructed by FIB tomography. Sample preparation, milling and WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


163

imaging conditions and data processing have been optimized to obtain a high quality representative 3D microstructure having the following dimension of 8.66×9.79× 11.41µm3. Home made programs have led to the calculation of some microstructural parameters such as volume fraction, connectivity, specific surface and interface areas and TPB length including the proportion of active and non active ones. These calculations have shown that for an anode with an initial 8YSZ/NiO volume ration of 60/40 leading after reduction to volume relative amounts 41% of pores, 33% of 8YSZ and 26% of Ni, all the phases were highly percolated. The active TPB length has been estimated as 7.4µm/µm3 which comparing to the literature results corresponds to an effective anode. By solving the diffusive transport equation with finite difference calculations, the effective thermal, electronic and ionic conductivities could be determined.

References [1] S. Singhal, K. Kendall, High-Temperature Solid Oxide Fuel Cells: Fundamentals, Design and Applications, Elsevier Advanced Technology, Oxford, UK, 2003. [2] P.R. Shearing, J. Gelb, N.P. Brandon, X-ray nano computerised tomography of SOFC electrodes using a focused ion beam samplepreparation technique, J. of the Europ. Ceram. Soc., 30 (2010) 1809-1814. [3] K.N. Grew, A.A. Peracchio, W.K.S. Chiu, Characterization and analysis methods for the examination of the heterogeneous solid oxide fuel cell electrode microstructure, Part 2 : Quantitative measurement of the microstructure and contributions to transport losses, J. Power Sources, 195 (24) (2010) 7943-7958. [4] J.R. Wilson, W. Kobsiriphat, R. Mendoza, H.-Y. Chen, J.M. Hiller, D.J. Miller, K. Thornton, P.W. Voorhees, S.B. Alder, S.A. Barnett, Threedimensional reconstruction of a solid-oxide fuel-cell anode, Nature Materials, 5 (2006) 541-544. [5] S.A.Barnett, J.R. Wilson, W. Kobsiriphat, H.-Y. Chen, R. Mendoza, J.M. Hiller, D.J. Miller, K. Thornton, P.W. Voorhees, S.B. Alder, Three dimensional analysis of solid oxide fuel cells using Focused ion beamScanning electron microscopy, Microsc. Microanal., 13 (2007) 596-597. [6] D. Gostovic, J.R. Smith, D.P. Kundinger, K.S. Jones, E.D. Washman, Three-dimensional reconstruction of porous LSCF cathodes, Electrochem. Solid-State Lett., 10 (12) (2007) B214-B217. [7] L. Holzer, B. Munch, B. Iwanschitz, M. Cantoni, T. Hocker, Th. Graule, Quantitative relationships between composition, particle size, triple phase boundary length and surface area in Ni-cermet for Solid Oxide Fuel Cells, J. Power Sources, doi : 10.1016/j.jpowsour.2010.08.006. [8] L. Holzer, B. Muench, M. Wegmann, P.H. Gasser, R.J. Flatt, FIB-nanotomography of particulate systems – Part 1 : Particle shape and topology of interfaces, J. Am. Ceram. Soc., 89 (2006) 2577-2585. [9] J.R Wilson, J.S Cronin, S.A. Barnett, Linking the microstructure, performance, and durability of Ni-Yttria-Stabilized Zirconia solid oxide WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


[10]

[11] [12]

[13]

[14]

[15] [16] [17] [18]

[19]

[20] [21]

fuel cell anodes using three dimensional focused ion beam – Scanning electron microscopy imaging, Scripta Materialia, doi 10.1016/j.scriptamat.2010.09.025. B. Ruger, J. Joos, A. Weber, T. Carraro, E. Ivers-Tiffée, 3D-modeling and performance evaluation of mixed conducting (MIEC) cathodes, ECS Transactions, 25 (2) (2009) 1211-1220. J. Hoshen, R. Kopelman, Percolation and cluster distribution. I. Cluster multiple labelling technique and critical concentration algorithm, Phys. Rev. B., 1 (14) (1976) 3438-3445. Y. Nakashima, S. Kamiya, Mathematica programs for the analysis of three-dimensional pore connectivity and anisotropic tortuosity of porous rocks using X-ray computed tomography image data, J. Nucl. Sci. Technol., 44 (9) (2007) 1233-1247. D. Rochais, G. Le Meur, V. Basini, G. Domingues, Microscopic thermal characterization of HTR particle layers, Nucl. Eng. And Design, 238 (2008) 3047-3059. T.G. Sherwood, R.L Pigford, C.R. Wilke, Mass transfer, In:Clark, B.J. Maisel, J.W. (Eds.) McGraw-Hill Inc., New York, pp. 39-43, 1975. S. Kakaç, A. Pramuanjaroenkij, X.Y. Zhou, A review of numerical modelling of solid oxide fuel cells, J. Hydrogen Energy, 32 (2007) 761-786. D.R. Lide, Handbook of Chemistry and Physics, Thermal Conductivity of gases, chap. 6, p 206, 89 th. ed., Taylor and Francis, Boca Raton (2008). K.W. Schlichting, N.P. Padture, P.G. Klemens, Thermal conductivity of dense and porous yttria-stabilized zirconia, J. Mat. Sci., 36 (2001) 3003-3010. Y.S. Touloukian, R.W. Powell, C.Y.Ho, P.G. Klemens, Thermophysical Properties of matter, vol. 1, Thermal conductivity, metallic elements and alloys., IFI/Plenum, New-York-Washington, 1970. R. Landauer, The electrical resistance of binary metallic mixtures, J. Appl. Phys., 23 (1952) 779-784. Z. Wu, M. Liu, Modelling of ambipolar transport properties of composite mixed ionic-electronic conductors, Solid State Ionics, 93 (1997) 65-84.



165

Modelling of load transfer between porous matrix and short fibres in ceramic matrix composites J. G. P. Silva1,2, D. Hotza1, R. Janssen2 & H. A. Al-Qureshi1 1 2

Federal University of Santa Catarina (UFSC), Florianópolis, SC, Brazil Technical University of Hamburg-Harburg(TUHH), Germany

Abstract The aim of this work is to present a model of load transfer between porous matrix and short fibres in ceramic matrix composites. This analysis is based on the earlier shear-lag models used for polymeric composites. However, geometry and strength of fibres in addition to the matrix porosity are included in the present analysis. The theoretical curves for the longitudinal and shear stresses distributions along the fibre-porous matrix interface are presented. They exhibited a maximum strength point at the middle of the short fibres. It became evident that the critical length is governed by the relative properties of the fibres, matrix and porosity, which greatly influenced the load carrying capacity of the fibres in the composites. In addition, the present simplified solution facilitates the understanding of the interface mechanism using porous matrix. Keywords: modelling, ceramic matrix composites, shear-lag models, porous ceramics.

1 Introduction Modern structural ceramic composites possess a number of unique properties that cannot be achieved by other materials. Therefore, they have a potential for saving energy, reducing wear, and increasing the lifetime of components [1]. However, regardless of their remarkable properties, structural ceramics are not as widely used in industry as they should and could be. Among the reasons for the reluctance of industry to introduce structural ceramic as components are [1]: • high price of ceramic parts, • insufficient knowledge in “traditional” metal-oriented engineering, WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line) doi:10.2495/MC110151

166 Materials Characterisation V • low toughness, • need of redesigning components to meet specific requirements. In the case of ceramic composites, the price factor is still the major limitation in terms of commercial applications. This is especially true for components made of composites with complex structures and requiring mechanical performance and reliability. To this aim, fibre-reinforced ceramic composites have been developed to overcome the fragility and unreliability of monolithic ceramics. Their main advantages are high-temperature resistance, low density, better corrosion resistance and adequate damage tolerance [2]. There is a wide range of fibre-reinforced ceramic composites depending on the chemical composition of the matrix and reinforcement, although currently only Cf-C/SiC composites produced by silicon infiltration have reached commercial production. However, fibre-reinforced ceramic composites based on oxides (alumina, mullite) can provide key benefits about long-term stability under oxidizing atmospheres [2]. Despite considerable interest in oxide ceramic composites in the past decade [3–5], there is still no production concept that meets the requirements in view of cost and performance. This study aims to investigate the load transfer phenomenon in ceramic-ceramic short fibre composites, using a simplified mathematical model that aims to predict the actual effectiveness of reinforcement depending on the material properties.

2 Modelling 2.1 Previous considerations and analysis Considering a loaded composite made of a dense fibre with length 2L embedded in a porous matrix made of the same material of the fibre, hereby is assumed that no slippage occurs between fibre and matrix. It should be also considered that the Poisson’s ratio of fibre and matrix is the same, which implies on the inexistence of transversal stress when the loading is applied along the fibre. Considering the displacements in the fibre u and far away from the fibre v: (Fig. 1):

Figure 1:

Simplified scheme of the stress field around the fibre. (a) without loading; (b) loaded [6].



167

From Hooke’s law:

  E 

P  E A L

(1)

Taking the differential:

dP EA d    dx L dx

(2)

Cox proposes a similar behaviour [7]:

dPf dx

 B (u  v )

(3)

where Pf is the load acting on the fibre and B is a constant that depends on the fibre distribution and the Young’s Modulus of fibre and matrix. Differentiation of Eq. (3) leads to:

d 2 Pf dx

2

 du dv   B    dx dx 

(4)

The derivatives of u and v can be taken as the deformations in the fibre and matrix, respectively:

Pf du  dx Af E f

(5)

dv  dx

(6)

Substitution of (5) and (6) in (4), gives:

d 2 Pf dx

2

 P   B f    Af E f   

(7)

A solution to this differential equation leads to:

Pf  E f A f   S sinh   x   T cosh   x  where:


(8)




B E f Af

(9)

and S and T are constants defined by the boundary conditions. To simplify the solution, instead of the stress distribution proposed by Cox, a model proposing that the stress drop in the extremity of the fibre follows a quadratic behaviour. To evaluate the model herein described, it is possible to apply the equations to a model composite, made of a porous alumina matrix and alumina fibres. Table 1 summarizes the properties that are relevant for the calculation. Table 1:

Properties of the model composite.

Property Fibre volume fraction Matrix porosity fraction Fibre length – 2l (mm) Fibre diameter (μm) Ratio Critical length/Length or α

Value 0.45 0.24 50.8 10 0.25

2.2 Proposed model According to Fig. 2, let us consider a composite with fibres whose length is 2L, diameter 2r and Young’s modulus Ef, embedded in a matrix with porosity ρ, made of the same material of the fibre. Hereby we define the critical length Lc, in which from the tip of the fibre the stress distribution isn’t constant by the shearlag between matrix and fibre. It is more feasible to work with α, the ratio between the critical length and fibre length, being Lc =α·L.

Figure 2:

Proposed stress distribution and boundary conditions.



169

It can be proposed that the stress distribution between the points L-αL and L follows a behaviour such as:

 f ( x )  Ax  B

(10)

By using the boundary conditions given in Fig. 2, and substituting them in (10):

E f   A( L   L)  B

(11)

(1   ) E f   AL  B

(12)

Isolating B in (12) and replacing in (11):

(1   ) E f   AL  E f   A( L   L)

(13)

  E f   A L

(14)

And then:

A

Ef L

(15)

By replacing A from (12):

(1   ) E f   

Ef B 

(16)

Therefore, B is given by:

1    B  1    1    E f     

(17)

By replacing the constants in (10), we have the stress distribution behaviour:

 f ( x)  

Ef 1    x  1    1    E f  L    

(18)

To determine the shear stresses along the fibre, the force equilibrium in a fibre element with diameter 2r and length dx is made in the x direction, resulting in: (19)    r 2    2 rdx  0 Then, the shear stresses are given by:

 

r d 2 dx


(20)

170 Materials Characterisation V By the differential of (18):

 f ( x) 

Figure 3:

r E f  2 L

(21)

Stress distribution along the fibre for different matrix porosities. α=0.25.

Figure 4:

Force equilibrium in a fibre element whose length is dx.

With the stress distribution along the fibre, it is possible to calculate the average stress carried by the fibre in the composite, given by:

f 

1 L  f ( x )dx L 0

(22)

For α ≥ 1, i.e. the fibre is shorter than the critical length:

f 

1 L Ef 1     x  1    1    E f  dx  0 L L    


(23)


Figure 5:

171

Shear stress distribution along the fibre for different matrix porosities. α=0.25.

Then,

   Ef  f  1    2  

(24)

And for 0 < α < 1, i.e., the fibre is longer than the critical length:

f 



L  L

0

E f  dx  

L

L  L

 f ( x )dx

(25)

L

Therefore:

 

 f  E f  1 

   2 

(26)

With the average stresses well defined, we can define the stresses in the ply longitudinal and transversal directions. When the matrix material is the same as the fibre, it is possible to write the elastic modulus of the matrix as a function of the fibre modulus:

Em  E f e  bp where b is a shape factor that depends on the pore shape and distribution.


(27)


Figure 6:

Longitudinal resistance of the composite [6].

The stress on the transversal direction is proportional to the matrix maximum stress, given by:

 T   m  E f  e bp

(28)

The stress on the longitudinal direction is given by the average value between matrix and fibre, based on the volumetric fractions of fibre and matrix:

 L   m 1   f    f  f  1   f  E f  e  bp   f  f

(29)

Therefore for 0 < α < 1:

 L  1   f  E f  e  bp  E f   1  

   f 2 

(30)

And for α > 1:

 

 L  1   f  E f  e  bp  E f  1   


  f 2 

(31)


Longitudinal ply ultimate strength, for different α.

Figure 7:

Figure 8:

173

Longitudinal ply ultimate strength for different porosities.

3 Conclusions A load transfer model was proposed to ceramic composites, which relates the matrix porosity and fibre length with the mechanical strength of such composites. It is hoped that the experimental validation of the model can be established or any changes aimed at increasing its accuracy can be made.



Acknowledgements The authors would like to thank the agencies CAPES (Coordenação de Aperfeiçoamento de Pessoal de Nível Superior), CNPq (Conselho Nacional de Desenvolvimento Científico e Tecnológico) and DFG (Deutsche Forschungsgemeinschaft), for funding this research under grant Bragecrim 015/09.

References [1] Janssen, R., Scheppokat, S. & Claussen, N., Tailor-made ceramic-based components – Advantages by reactive processing and advanced shaping techniques. Journal of the European Ceramic Society, 28, pp. 1369-1379, 2008. [2] Janssen, R. & Hotza, D. Low-cost and reliable production of oxide ceramic matrix composites, Bragecrim Project Proposal, December 2008. [3] Wendorff, J., Janssen, R. & Claussen, N., Saphirfaserverstärkung reaktionsgebundener oxider Keramiken. Verbundwerkstoffe und Werkstoffverbunde, ed. G. Ziegler, pp. 421-424, 1996. [4] Lundberg, R. & Eckerbom, L., Design and processing of Al-oxide composites. Ceramic Transactions, 58, pp. 95-104, 1995. [5] Levi, C.G., Yang, J.Y., Dalgleish, B.J., Zok, F.W. & Evans, A.G., Processing and performance of an all-oxide ceramic composite. Journal of the American Ceramic Society, 81, pp. 2077-2086, 1998. [6] Casaril, A.; Gomes, E.R.; Soares, M.R.; Fredel, M.C. & Al-Qureshi, H.A., Análise micromecânica dos compósitos com fibras curtas e partículas. Matéria, 12(2), pp. 408-419, 2007. [7] Cox, H.L., The elasticity and strength of paper and other fibrous materials. British Journal of Applied Physics, 3, pp. 72-79, 1952



175

Modeling aspects concerning the axial behavior of RC columns H. O. Koksal1, T. Turgay 2, C. Karakoç3 & S. Ayçenk4 1

Construction Technology Program, Çanakkale Onsekiz Mart University, Turkey 2 Faculty of Engineering and Architecture, Abant İzzet Baysal University, Bolu, Turkey 3 Civil Engineering Department, Boğaziçi University, Turkey 4 M.S. Student, Civil Engineering Faculty, Yıldız Technical University, Turkey

Abstract This paper is concerned with the axial behavior of the RC columns. Stress-strain relationships of experimentally tested RC columns under concentric loading are compared with the predictions of the Koksal model. Moment-curvature analyses of RC sections are also performed employing the same model in a self-developed moment-curvature program for confined concrete. Results are compared with the output of EXTRACT which uses the Mander concrete model. Keywords: confinement, reinforced concrete columns, concentric loading, stress–strain relations.

1 Introduction Confining pressure on RC columns increases the concrete strength. Transverse reinforcements such as steel stirrups, FRP wraps, and steel jackets create a triaxial compressive stress state producing confinement action around the concrete core. Extensive research on the improvement of concrete confinement has been carried out since the pioneering study of Richart et al. [1]. The constitutive model for confined concrete plays an important role on the accuracy of the moment-curvature curves to perform a reliable pushover analysis of RC frames. There are some frequently cited models (Hognestad [2], Kent and Park [3], Sheikh and Uzumeri [4], Mander et al. [5], Saatcioglu and Razvi [6]) to predict WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line) doi:10.2495/MC110161

176 Materials Characterisation V the peak stress or the stress–strain curve of confined concrete. Köksal has also proposed a failure criterion for concrete under tri-axial compression stress state [7]. In this paper, four square RC columns tested in Yıldız Technical University (Turgay [8], Köksal et al. [9]) are evaluated implementing Koksal model. The flexural behavior of the columns is also simulated and their moment-curvature relations corresponding to different axial load levels are obtained using the same model. Moment-curvature curves are compared with the outputs of EXTRACT.

2 Experimental study The square specimens in the test program have 200x200 mm cross-section dimensions and 1000 mm height. The columns were tested in the structural laboratory of Yıldız Technical University [8]. Figure 1 shows the details of test setup and instrumentation for the two specimens.

Figure 1:

Test setup and details of test specimens C1L4S8 and C1L8S8 [8].

The experimental study was limited to one type of concrete mix design typically. C1 type columns are tested at 30 days. All longitudinal bars are 10mm in diameter and L4 and L8 shows the number of the bars in the cross-section of a column. The tie spacing is 100mm and the S8 represents diameter of stirrups. For measurement of axial strains, four linear variable displacement transducers (LVDTs) are placed over the central 400mm gage length at each side of a column in a similar way used to assess any eccentricity of the applied load as recommended in the study of Shrive et al. [10]. Although the LVDT readings were provided very close to each other, next to the maximum axial load there can be a significant variation between the minimum and maximum values of shortening reaching very high values [9]. Figure 2 shows the axial load-axial WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


177

shortening curve for C1L8S8 for which almost perfectly homogeneous strains were obtained. However, for other columns the differences between LVDT readings are somewhat greater indicating a localized damage zone at one specific region of the specimen.

Figure 2:

Four separate LVDT readings of axial load-shortening curves for C1L8S8 [9].

3 Simulation of flexural behavior and moment curvature relations The limit state design procedure of reinforced concrete elements has undergone major revision by most of the international codes in harmony with the performance-based design engineering approach. In this approach, momentcurvature behavior of flexural members is needed to define the deflection demand and to simulate the behavior of the reinforced concrete members under lateral loads such as earthquake actions. Moment-curvature plots readily illustrate stiffness, strength, and cross-sectional ductility, and allow the calculation of deflections after materials become nonlinear. In a flexural member the shear reinforcement or any other confinement mechanism applies pressure to the concrete in the compression zone and affects energy dissipation capacity directly. Therefore, to predict the moment-curvature behavior of a flexural member, the stress–strain behavior of confined concrete under axial compression is vital. With the development of performance-based design methods, there is an increasing need for simplified but reliable analytical models capable of predicting the flexural behavior of reinforced concrete members. The flexural behavior of reinforced confined concrete sections is introduced by Koksal model [7]. The failure criterion proposed by Koksal is given as:

f 1.1317 0.7645    2 k1, 2 ,3   0 WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)

(1)






where k  k f l , f c' ,  and  are deviatoric and hydrostatic lengths, respectively. In this equation k, as a function of the lateral confinement pressure, f l , and the cylinder compressive strength of concrete, f c' can be expressed as [7] : 2    fl  fl  k  4.07 '  0.89 '   0.807 fc'   fc  fc   

(2)

The theoretical confining pressure, f l , can be found in usual way as [11] : or (3) f l , x  k e  sx f yh f l , y  k e  sy f yh In this relation,  sx and  sy are the confinement proportion both x and y direction respectively, ke is the confinement effectiveness coefficient, and fyh is the yield strength of transverse reinforcing steel. The Saenz’s equation [12] is adopted for describing the monotonic stress-strain relationship for confined concrete:

1 

1E0

     E 1   0  2  1    1   cc    cc  Es

  

2

(4)

where  1 and 1 are axial compressive stress and strain of concrete, respectively; E0 is the initial tangent modulus of elasticity in MPa and can be calculated as E0  4750 f co' ; Es is the secant modulus at the point of maximum compressive stress f cc' which can be determined using Equation 1. The strain cc corresponding to the maximum compressive stress f cc' can be found employing the recommended relations of Richart et al. [1] :



 cc   ' c 1  k 2 

fl f 'c

  

(5)

where  c' is the peak strain at the strength of plain concrete cylinders. In this equation, k2 is taken as 5k1. Figure 4 aims to compare experimental axial stress-strain curves of the four column specimens with those predicted from Eqs (1)-(4). In this study, a simple program is written in FORTRAN to produce momentcurvature plots for confined column section under bending and axial compression. The cross section is divided into 1mm thickness slices. For a given depth of natural axis, the strain at the extreme slice in compression is found by iterative procedure. For each slices, the stress and strain are calculated using force equilibrium and compatibility requirements respectively. Stress acting on the core concrete (cover concrete is neglected) calculated using Koksal’s model. Koksal’s model has also been utilized through the 3D finite element analysis of RC and FRP-confined concrete columns [11, 13] successively. Figure 3(a) WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


179

demonstrates the typical axial stress-strain curves recommended by Saenz for a concrete member [12]. Finite concrete forces for unconfined and confined core of each slice are found by multiplying stress by corresponding areas. Stress at the reinforcement bars is found by entering a simple bi-linear stress-strain curve in Figure 3(b) with the strain value found from compatibility requirements at each load level. Finite forces on steel can be found by multiplying stress by the area of the reinforcement.

Figure 3:

(a) Saenz curve for uniaxial behavior of concrete; (b) stress-strain relation for steel.

Figure 4:

Experimental and predicted stress-strain curves for the specimens.


180 Materials Characterisation V Figure 5 show the moment-curvature relations for C1L4S8 and C1L8S8 at various load levels. Increasing the axial load from 200kN to 400kN always results an increase in the moment carrying capacity. But when a load near to the axial load capacity is considered, moment carrying capacity does not increase any more. Also decrease in the curvature is very obvious when the axial load increases as expected.

Figure 5:

Moment-curvature diagrams for the four columns at various axial load levels.

30

Moment (KNm)

C1L4S8

20

10 Model L200 EXTRACT L200 0 0

Figure 6:

20

40 60 Curvature x 10e6 (rad/m)

80

100

Output of EXTRACT program and the proposed model for moment-curvature diagrams for C1L4S8 at the axial load level of 200kN.



181

Figures 6–8 show the output of EXTRACT program [14] for C1L4S8 subjected to the axial load levels of 200kN, 400kN and 600kN respectively. EXTRACT program employs the Mander model for confined concrete. There are not significant differences if one compares the curvature values for both programs. 40 C1L4S8

Moment (KNm)

30

20

10 Model L400 EXTRACT L400 0 0

20 40 Curvature x 10e6 (rad/m)

60

Figure 7:


Figure 8:




4 Conclusion This paper has been primarily concerned with the modeling aspects for the axial behavior of RC columns. Experimental axial stress-strain curves of the four column specimens are predicted employing Koksal model. Furthermore, a spreadsheet-based program is constructed to produce moment-curvature plots for the performance-based design engineering approach. The following conclusions can be drawn based on the results of the analyses: 1. As can be seen in Figure 5, Koksal model successively predicts the axial behavior of the four RC column specimens tested under concentric loading. 2. It can be easily observed that Koksal model results into somewhat higher values for the moment carrying capacities than the outputs of EXTRACT program which employs Mander model in Figure 6–8, but the general trend of the curves are the same. The developed program gives 27kNm and 38kNm for moment carrying capacities at the axial load levels of 400kN and 600kN while EXTRACT program results are between 22kNm and 27kNm.

Acknowledgement The support of Boğaziçi University Research Fund (ref: research project 5232) for this paper is gratefully acknowledged.

References [1] Richart, F.E., Bradtzaeg, A. and Brown, R. L. 1928. A study of the Failure of Concrete under Combined Compressive Stresses. Bulletin Np. 185, Engineering experimental station University of Illinois, Urbana, pp. 104. [2] Hognestad, E., 1951. A Study of Combined Bending and Axial Load in Reinforced Concrete Members. Bulletin Series No.399, University of Illinois Eng. Exp. Station, Urbana. [3] Kent, D.C. and Park, R. 1971. Flexural Members with Confined Concrete. Journal of the Structural Division, Proc. of the American Society of Civil Engineers, 97(ST7), pp.1969-1990. [4] Sheikh, S.A. and Uzumeri, S.M. 1982. Analytical Model for Concrete Confinement in Tied Columns. Journal of the Structural Division, Proc. of the American Society of Civil Engineers, 108(ST12), pp. 2703-2722. [5] Mander, J.B., Priestly, M.J.N. and Park, R. 1988. Theoretical Stress-Strain Model for Confined Concrete. Journal of the Structural Engineering, ASCE, 114(8), pp.1804-1826. [6] Saatcioglu, M. and Razvi, S.R. 1992. Strength and Ductility of Confined Concrete. Journal of the Structural Engineering, ASCE, 118(6), pp.1590-1607. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


183

[7] Köksal, H.O. 2006. A Failure Criterion for RC Members Under Triaxial Compression. Structural Engineering and Mechanics, Techno-Press, 24(2), pp.137-154. [8] Turgay, T. CFRP Uygulanmıs_ betonarme elemanların performansı (The performance of FRP strengthened structural members). PhD thesis, Submitted to Yıldız Technical University; 2007. [9] Köksal, H.O., Karakoç, C., Polat Z., Turgay T. and Akgün Ş. 2007.Evaluation of Experimental Procedures for Confined Concrete Columns. Computational Methods and Experimental Measurements XIII, WIT Transactions on Modeling and Simulation, 46, 233-242. [10] Shrive, P.L., Azarnejad, A., Tadros, G., McWhinnie, C. and Shrive, N.G. 2003. Strength of Concrete Columns with Carbon Fibre Reinforcement Wrap. Canadian Journal of Civil Engineering, 30, pp. 543-554. [11] Doran, B. 2009.Numerical simulation of conventional RC columns under concentric loading. Material and Design, 30(6), 2158-2166. [12] Saenz, L.P. 1964. Discussion of Equation for Stress-Strain Curve of Concrete by Desai and Krishnan. ACI, 61(9), 1229-1235. [13] Köksal, H.O, Doran, B., and Turgay, T. 2009. A practical approach for modeling FRP wrapped concrete columns. Constr. and Build. Mat., 23(3), 1429–1437. [14] EXTRACT-v.3.0.8, Cross section analysis program of structural engineers, TRC/Imbsen Software Systems.



Section 5 Innovative experiments



187

Surface characterization of eucalyptus and ash wood veneers by XPS, TOF-SIMS, optic profilometry and contact angle measurements G. Vázquez, R. Ríos, M. S. Freire, G. Antorrena & J. González-Álvarez Department of Chemical Engineering, University of Santiago de Compostela, Spain

Abstract Composition and properties of the wood surface are very important in board manufacture as they determine wood-adhesive bonding and the quality of the final product. In this work two spectroscopic techniques, X-ray photoelectron spectrometry (XPS) and time of flight secondary ion mass spectrometry (TOF – SIMS), have been employed to study the surface composition of eucalyptus (Eucalyptus globulus) and ash (Fraxinus excelsior) sliced cut veneers. Both wood species are widely used for decorative veneers in the finishing of wood panels (particle boards, medium density fiber–boards, etc.). Further characterization of the wood surface was carried out by optic profilometry and wettability analysis, using the sessile drop method to measure contact angles. Wood is a very heterogeneous material and its composition can vary significantly depending on the sampling area; then, the influence of the radial position of the veneer in the trunk was also analyzed. From the low resolution XPS spectra the oxygen to carbon (O/C) ratio was calculated and, resolving the C1s signal, the percentages of the different carbon peaks corresponding to different functional groups (C1, C2, C3 and C4 carbons) were also calculated. The lower O/C ratio and the higher C1/C2 ratio for eucalyptus than for ash was attributed to the higher concentration of extractives on the eucalyptus wood surface that was confirmed by the TOF-SIMS spectra, which additionally revealed a patchy distribution of the extractives. The higher hydrophobicity of the eucalyptus wood surfaces was also supported by the results on the wetting properties of the veneers. Eucalyptus veneers exhibited a significantly higher constant wetting rate angle (cwra) and, consequently, a lower wettability than WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line) doi:10.2495/MC110171

188 Materials Characterisation V ash veneers, which can also be related with the higher values of the rugosity parameters for the latter. There were no significant influences of the radial position of the veneer on the surface properties. Keywords: wood veneers, wettability, contact angles, XPS, TOF-SIMS, optic profilometry.

1 Introduction Wood is a renewable resource composed mainly of three polymers, cellulose, hemicelluloses and lignin, and various extractives such as fatty acids, sterols, tannins, etc. Wood surface characteristics condition the interaction with adhesives and binders in wood based products and, therefore, have influence in the quality of the final product. Several methods have been proposed to characterize material surfaces. Contact angle analysis is one the traditional methods for the characterization of surfaces on a macroscopic scale, and provides information on surface wetting properties [1]. Surface chemical composition of complex materials such as wood can be determined using two complementary spectroscopic techniques, X-ray photoelectron spectroscopy (XPS) and time-of-flight secondary ion mass spectrometry (TOF-SIMS), the former generating more quantitative data and the latter more qualitative data [2]. X-ray photoelectron spectroscopy (XPS) provides the elemental composition for the outermost 5–10 nm surface layers and functional groups, chemical bonding types, oxidation state, etc. can be additionally deduced from XPS data. Time-of-flight secondary ion mass spectrometry (TOF-SIMS) provides data for depths of 1–2 nm, including detailed chemical information via identification of intact molecular ions or characteristic molecular fragments that are emitted from the surface [3]. Additionally, it provides ion images, mapping the lateral distribution of secondary ions signals from lignin, carbohydrates, extractives and metals within an analyzed wood surface area [2, 4]. The aim of this paper was to evaluate and compare the surface properties and composition of eucalyptus (Eucalyptus globulus) and ash (Fraxinus excelsior) sliced cut veneers used as decorative veneers in the finishing of wood panels. A wettability analysis, using the sessile drop method to measure contact angles, was applied in combination with two spectroscopic techniques, X-ray photoelectron spectrometry (XPS) and time of flight secondary ion mass spectrometry (TOF –SIMS). Roughness of the wood surface was analysed by optic profilometry. Wood is a very heterogeneous material and its composition can vary significantly depending on the sampling area; then, the influence of the radial position of the veneer in the trunk was also analyzed.

2 Methods 2.1 Contact angle determination Samples of eucalyptus (Eucalyptus globulus) and ash (Fraxinus excelsior) sliced cut veneers with a humidity of 8% (wet basis) were used to prepare 2 cm (length) WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


189

x 2cm (width) x 0.6 mm (thickness) wood pieces. Contact angle measurements were made perpendicular to the grain direction of the wood according to the sessile drop method using a Dataphysics OCA 15 Plus equipment (Filderstadt, Germany) with a video measuring system with a high-resolution CCD camera. The data were analyzed with the Dataphysics software SCA 20. To compare the wettability of both wood species and to analyze the influence of the radial position of the veneer in the trunk, 20% (by weight) aqueous solutions of chestnut shell extracts, potential components of wood adhesives [5], were used. Ten μL drops were added and the variation of contact angle with time was registered till 240 s. Contact angles were calculated according to the Young-Laplace method. The constant wetting rate angle (cwra) was determined as per Nussbaum [6] plotting the wetting rate (dθ/dt, being θ the contact angle) against time and selecting the θ value (cwra) corresponding to a constant wetting rate. Initial (θ0, t=0), final (θf, t=240 s) and cwra contact angles were calculated as the average of thirty measurements on each wood species and radial position. 2.2 X-ray photoelectron spectrometry (XPS) Analysis of the samples was performed using a Thermo Scientific K-Alpha ESCA instrument equipped with aluminium Ka1,2 monochromatized radiation at 1486.6 eV X-ray source. Due to the non conductor nature of samples it was necessary to use an electron flood gun to minimize surface charging. Neutralization of the surface charge was performed by using both a low energy flood gun (electrons in the range 0 to 14 eV) and an electrically grounded stain steel screen placed directly on the sample surface. The XPS measurements were carried out using monochromatic Al-K radiation (hν=1486.6 eV). Photoelectrons were collected from a takeoff angle of 90º relative to the sample surface. Measurements were done in a Constant Analyser Energy mode (CAE) with a 100 eV pass energy for survey spectra and 20eV pass energy for high resolution spectra. Charge referencing was done by setting the lower binding energy C1s photopeak at 285.0 eV C1s hydrocarbon peak. The spectra fitting was based on Chi-squared algorithm used to determine the goodness of a peak fit. Surface elemental composition was determined using the standard Scofield photoemission cross sections. The chemical functional groups identity was obtained from the high-resolution peak analysis of carbon-1s (C1s) and oxygen1s (O1s) envelopes. 2.3 Time of flight secondary ion mass spectrometry (TOF –SIMS) The TOF-SIMS analyses were made on a TOF-SIMS IV (ION-TOF GmbH, Germany).The sample was bombarded with a pulsed Bismuth ion beam. The secondary ions generated were extracted with a 10 kV voltage and their time of flight from the sample to the detector was measured in a reflectron mass spectrometer. Mass spectra of positive and negative secondary ions were acquired from randomly selected areas of 500 x 500 µm2 on each sample. In all WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)

190 Materials Characterisation V cases, a 25 keV pulsed Bi3+ beam at 45º incidence was used. Low–energy electron flooding was used for charge compensation. All analyses were done under static SIMS conditions (1012 ions/cm3). 2.4 Optic profilometry Interferometric profilometry measurements were carried out in the Vertical Scanning Interferometry (VSI) mode using an Interferometric Microscope WYKO NT-1100. For each sample, three measurements were carried out at a 5X magnification (1.2x0.9 mm2), to determine statistically representative values of surface roughness parameters, and a measurement at magnifications of 20X (298x227 µm2) and 50X (119x91 µm2) to see the details of the surface topography. Great field measurements covering an area of 3x2 mm2 at a 5X magnification were also performed to obtain a more representative image of the surface.

3 Results and discussion 3.1 Wettability of eucalyptus and ash wood veneers Figure 1 shows the evolution of contact angle with time for an aqueous solution of chestnut shell extracts on eucalyptus and ash veneers. For each species, veneers obtained from different positions in the trunk, interior and exterior, were compared. Throughout the time range tested, contact angles for eucalyptus veneers were significantly higher than those for ash veneers. Additionally, for both wood species there were not significant differences between the contact angles for the interior and exterior veneers.

Figure 1:

Contact angle versus time for an aqueous solution of chestnut shell extracts on interior and exterior eucalyptus and ash sliced veneers.



191

The values calculated for the initial (t=0; θ0), final (t =240 s; θf) and constant wetting rate (cwra) angles are presented in Table 1. Not only initial contact angles but also final and constant wetting rate angles were remarkably higher for the eucalyptus veneers. Higher contact angles are indicative of a worse wetting behaviour and could be related with a greater migration of hydrophobic substances to the surface of eucalyptus veneers [7]. Differences between the θ0, θf and cwra values of interior and exterior veneers were almost negligible which confirms the previous qualitative observation. Table 1:

Initial (θ0), final (θf) and constant wetting rate (cwra) angles for exterior and interior veneers of eucalyptus and ash wood with an aqueous solution of chestnut shell extracts.

Veneer Eucalyptus Ash

θ0 (º) 99.72 (5.83) 75.22 (1.44)

Exterior θf (º) cwra (º) 45.25 50.8 (1.88) (0.11) 28.01 27.99 (0.37) (0.37)

θ0 (º) 95.34 (4.77) 73.42 (9.91)

Interior θf (º) cwra (º) 46.70 50.12 (6.77) (0.10) 27.75 30.6 (8.43) (0.76)

(Standard deviation)

3.2 X-ray photoelectron spectrometry (XPS) Figure 2a shows the XPS survey spectra of eucalyptus and ash samples of interior veneers. The surface elemental compositions (in atomic %) for the different sample groups are summarized in Table 2. The percentages shown are the mean value of three determinations. Table 2:

Surface elemental composition (in atomic %) for eucalyptus and ash veneers.

Veneer Exterior eucalyptus Interior eucalyptus Exterior ash Interior ash

C 80.7 (2.2) 79.2 (0.1) 77.2 (1.0) 76.9 (1.4)

O 18.0 (2.3) 19.3 (0.7) 20.8 (1.1) 20.5 (1.2)

N 0.7 (0.2) 1.0 (0.4) 0.6 (0.1) 0.9 (0.1)

Si 0.6 (0.1) 0.8 (0.2) 1.2 (0.5) 1.2 (0.5)

Ca 0.3 (0.1) 0.5 (0.1)


As seen in Table 2, carbon and oxygen are the main components in the wood surfaces and small amounts of nitrogen and silicon were also detected in both wood species, whereas calcium was present only in the ash samples. Nitrogen is present in the form of amine/amide functional groups. This assignment is as a result of the N1s peak position, which is at approximately 400 WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)

192 Materials Characterisation V eV (with C-C carbon set at 285.0 eV as the binding energy scale reference). The presence of such compounds is mainly as a result of naturally occurring proteins in wood but some contamination by adsorption from the laboratory air cannot be excluded [8]. From the high resolution spectra (Figure 2b) the functional groups which made up the C1s spectra have been identified. The C1s spectra of all samples comprised four peaks with binding energies (BE) of approximately 285.0 (C1), 286.6 (C2), 288.1 (C3) and 289.2 (C4) eV. The C1 carbon component is related to C–C or C–H bonds, the C2 carbon component represents single C–O bonds, the C3 carbon component is bonded to a carbonyl or two non-carbonyl oxygen (C=O or O-C-O) and the C4 class of carbon atoms is bonded to a carbonyl and a non-carbonyl oxygen (O=C-O) [9].

Figure 2:

XPS survey spectra (a) and high resolution C1s spectra (b) for interior eucalyptus and ash veneers.

The relative percentages of the C1s components are shown in Table 3 together with the O/C and C1/C2 ratios. The C1 component arises from lignin and extractives and the C2 component can arise from all wood components but predominantly from cellulose. For both species, most carbon atoms (60–66.5%) present in the wood surface layer were C1-carbons, and between 29.2 and 33.5% of the total carbon signal was from the C2 type. According to a normal wood WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


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composition and the structure of wood components, carbon atoms bonded with one oxygen atom (C2 component) should be greater. The observed increase in the presence of C1 components, which was more significant for eucalyptus samples, might be due to the migration of lipophilic extractives to the wood surface [9]. Table 3:

Percentages of the C1s spectra components and C1/C2 and O/C ratios for eucalyptus and ash veneers.

Veneer Exterior eucalyptus Interior eucalyptus Exterior ash Interior ash

C1 66.5 (4.7) 65.4 (2.9) 60.1 (1.7) 60.0 (2.2)

C2 29.2 (3.4) 29.3 (3.2) 32.4 (0.4) 33.5 (2.4)

C3 2.3 (1.6) 2.8 (0.5) 5.0 (1.2) 4.1 (1.0)

C4 2.0 (0.4) 2.8 (0.3) 2.6 (0.1) 2.4 (0.7)

C1/C2

O/C

2.3

0.22

2.2

0.24

1.9

0.27

1.8

0.27


This theory is reinforced by the values obtained for the O/C and C1/C2 ratios, which are directly related to the chemical composition of wood constituents (polysaccharides, lignin and extractives). The O/C ratio varied between 0.22 and 0.27 (Table 3). If experimental O/C ratio values are compared to those of pure compounds, 0.83 for cellulose, 0.33 for lignin and 0.1 for extractives [10] the preferential presence of lignin and extractives on the wood surface is demonstrated, mainly in the case of eucalyptus veneers. Moreover, the C1/C2 ratio provides an additional evidence to support the above interpretations. Theoretically calculated values for the C1/C2 ratio are equal to 0 for pure cellulose, around 1 for lignin and 10 or higher for extractives [11]. Therefore, values in the range from 1.8 to 2.3 confirm the presence of extractives on the wood surfaces and especially on eucalyptus wood. Finally, when comparing the mean values of the C1/C2 and O/C ratios and the C1-carbon percentages, more hydrophobic material was found in exterior eucalyptus veneers. The more hydrophobic character of the eucalyptus wood samples is in agreement with the results obtained from contact angle measurements which revealed better wetting properties for ash veneers. 3.3 Time of flight secondary ion mass spectrometry (TOF–SIMS) TOF-SIMS spectra were prepared to represent the masses from 1 to 700 u, although the range from 100 to 200 u is the most interesting because in this spectral region fragments of lignin and carbohydrates are detected. On the contrary, in the region below 100 u, multiple fragments which are common to all organic materials, the so-called non-specific organic fragments appear and, therefore, fragments of this region of the spectrum are rarely used in the study of WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)

194 Materials Characterisation V organic compounds. Na, Ca or Al, are also detected in the area of the spectrum below 100 u. However, the intensity of these ions is low when compared with the intensity of the peaks of hydrocarbons. Extractive organic compounds appear in areas of the spectrum with m/z values between 230 and 700 u. Figure 3 shows the comparative positive polarity spectra for ash and eucalyptus veneers at m/z values from 135 to 185 u. Negative polarity spectra were nor included due to the low intensity of the observed signals. For both samples, the characteristic ions of lignin, at m/z =137 and 151 for lignin guaiacyl units (G), and at m/z=167 and 181 for lignin syringyl units (S), were observed.

Figure 3:

Comparison of the positive ion TOF-SIMS spectra from 135 to 185 u for eucalyptus and ash veneers.

Typical peaks of cellulose and hemicelluloses, localized at m/z = 115, 127, 133 and 145 u, also appeared. Peaks at m/z =115 and 133 u are attributed to xylan, the peak at 127 u to mannan and the one at 145 u to cellulose. Additional peaks at m/z =147 and 149 u are attributed to mannan and cellulose. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


195

TOF-SIMS technique is not quantitative and, therefore, direct comparison of the intensity of a particular ion in different samples does not allow extracting conclusions about its abundance. However, it is possible to compare different samples by referring the intensities of the ions of interest to that of another ion not related to them. In this way, the intensities of the lignin ions at m/z = 137, 151, 167 and 181 u in eucalyptus and ash wood samples were referred to that of the ion at m/z =147 u, attributable to hemicelluloses and present in both kind of samples. The relative intensity of lignin ions in eucalyptus samples was slightly higher than in the ash ones, which indicated that lignin was more abundant on the eucalyptus veneer surfaces. Figure 4 presents the comparison of the positive polarity TOF-SIMS spectra for eucalyptus and ash veneers in the region from 135 to 450 u. Signals at 383, 397 and 425 u are attributed to organic wood extractives. As shown in the spectra, the content of extractives is higher in the eucalyptus sample than in the ash one which is consistent with XPS results.

Figure 4:

Comparison of the positive ion TOF-SIMS spectra from 135 to 450 u for eucalyptus and ash veneers.

Positive ion images for selected ions signals from eucalyptus and ash samples are shown in Figure 5. Black colour indicates no ion signal and white colour maximum signal intensity. Images of all the characteristic ions for lignin (137, 151,167 and 181 u) and extractives, mainly sterols (383, 397 and 425 u), are presented together. Ion images showed that lignin and mainly extractives were more prominent on eucalyptus than on ash veneer surfaces and, moreover, revealed that both components were inhomogeneously distributed on the surface.



Figure 5:

Positive ion images for eucalyptus (a) and ash (b) samples.

3.4 Optic profilometry Table 4 shows the average values of the roughness parameters Ra (arithmetic mean of the absolute value of the distances from the mean line to the profile) and Rq (mean value of the square roots of the distances from the mean line to the profile) for eucalyptus and ash veneer samples. Due to the heterogeneity of wood samples, surface roughness values depended strongly on the topographic characteristics of each particular area, so that the values showed a high dispersion. Therefore, the most representative roughness values might be those of the larger areas (called stitch), as the measure itself represents the average of a larger area of the samples. The main conclusion drawn from the values of surface roughness is that ash samples are rougher than eucalyptus samples, approximately 1.7 times based on the results of the stitch areas, which can be related with the better wettability behaviour previously encountered.



Table 4:

197

Roughness average values for eucalyptus and ash veneer surfaces.

Sample

Magnification 5X (1.2x0.9 mm2) 20X (298x227 µm2 )

Eucalyptus 50X (119x91 µm2) STITCH (3x2 mm2) 5X (1.2x0.9 mm2) 20X (298x227 µm2 ) Ash 50X (119x91 µm2) STITCH (3x2 mm2)

Ra (µm) 11.31 (2.44) 3.44 (1.24) 2.30 (0.91) 15.91 (0.45) 12.17 (1.23) 8.78 (2.60) 6.10 (2.56) 27.38 (2.55)

Rq (µm) 16.26 (4.63) 5.11 (0.66) 2.93 (1.00) 23.56 (4.70) 15.92 (1.23) 11.61 (4.14) 8.05 (3.54) 39.36 (6.80)


Acknowledgements This work was funded by Ministerio de Ciencia e Innovación, FEDER Funds and Plan E Fundy (CTQ2009-07539).

References [1] Bryne, L.E. & Walinder, M.E.P., Ageing of modified wood. Part 1: Wetting properties of acetylated, furfurylated, and thermally modified wood. Holzforschung, 64, pp. 295-304, 2010. [2] Englund F., Bryne L.E., Ernstsson M., Lausmaa J. & Walinder M., Spectroscopic studies of surface chemical composition and wettability of modified wood. Wood Mater. Sci. Eng., 1-2, pp. 80–85, 2009. [3] Bryne, L.O., Lausmaa, J., Ernstsson, M., Englund, F. & Walinder, M.E.P., Ageing of modified wood. Part 2: Determination of surface composition of acetylated, furfurylated, and thermally modified wood by XPS and ToFSIMS. Holzforschung, 64, pp. 305–313, 2010. [4] Tokareva, E.N., Fardim, P., Pranovich, A.V., Fagerholm, H.P., Daniel, G. & Holmbom, B., Imaging of wood tissue by ToF-SIMS: Critical evaluation and development of sample preparation techniques. App. Surf. Sci., 253, pp. 7569–7577, 2007. [5] Vázquez, G., González-Alvarez, J., Santos, J., Freire, M.S. & Antorrena, G., Evaluation of potential applications for chestnut (Castanea sativa) shell



[6]

[7]

[8] [9]

[10] [11]

and eucalyptus (Eucalyptus globulus) bark extracts. Ind. Crops Prod., 29, pp. 364-370, 2009. Nussbaum, R.M., Natural surface inactivation of Scots pine and Norway spruce evaluated by contact angle measurements. Holz Roh- Werkstoff, 57, pp. 419-424, 1999. Christiansen, A.W., Effect of ovendrying of yellow-poplar veneer on physical properties and bonding. Holz Roh- Werkstof , 52, pp. 139-149, 1994 Popescu, C.N, Tibirna, C.N. & Vasile C., XPS characterization of naturally aged wood, App. Surf. Sci., 256, pp. 1355–1360, 2009. Sinn, G., Reiterer, A. & Stanzl–Tschegg, S.E., Surface analysis of different wood species using X-ray photoelectron spectroscopy (XPS). J. Mater. Sci., 36, pp. 4673-4680, 2001. Barry, A.O., Koran, Z. & Kaliaguine, S., Surface analysis by ESCA of sulfite post-treated CTMP. J. Applied Polym. Sci., 39, pp. 31–42, 1990. Sernek, M., Kamke1, F.A. & Glasser, W.G., Comparative analysis of inactivated wood surface, Holzforschung, 58, pp. 22–31, 2004.



199

Interface resistances in heat and moisture transport: semi-scale experimental analysis Z. Pavlík, J. Mihulka, J. Žumár, M. Pavlíková & R. Černý Department of Materials Engineering and Chemistry, Faculty of Civil Engineering, Czech Technical University in Prague, Czech Republic

Abstract A sophisticated semi-scale system which allows experimental simulation of heat and moisture transport in building structures is employed for analyzing heat and moisture transfer across interfaces in stone masonry formed by argillite walling blocks and lime-pozzolana mortar. The dimensions of investigated structure are close to a real wall but the measuring accuracy is the same as in a laboratory experiment. Combined relative humidity and temperature mini-sensors are utilized in continuous long-term monitoring. Measured temperature and relative humidity profiles provide information on the properties of the interface between the walling material and mortar which makes possible to identify the interface permeability. This parameter can then be used in computational models of heat and moisture transport in masonry which adds to the accuracy of model predictions. Keywords: heat and moisture transport, interface resistances, semi-scale experiment, stone masonry, argillite, lime-pozzolana mortar.

1 Introduction Computational modelling of coupled moisture and heat transport represents an effective tool for prediction of behaviour of building structures exposed to climatic load. For reliability of computational analysis, two requirements must be met. At first, mathematical-physical model describing the coupled moisture and heat transport must be formulated. From the theoretical point of view, this model should be sufficiently sophisticated and complex, and should take into account all the external and internal effects that affect the moisture and heat WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line) doi:10.2495/MC110181

200 Materials Characterisation V transport. Typical example of such complex models represents for instance hybrid model developed by Grunewald [1, 2] that takes into account both diffusion and convection effects on moisture and heat transport. Although these models describe the physical reality of coupled moisture and transport in sophisticated way, their practical application for hygrothermal analysis of building structures is often limited by the accuracy and availability of input parameters that represent the second limiting requirement of computational modelling. There are two types of input parameters which have to be known in advance for precise computational analysis. The first are initial and boundary conditions. In case of analysis of existing structure, for example by reason of intended reconstruction, initial conditions can be determined using on site analysis of moisture and temperature fields in the studied structure. Boundary conditions are of two types. The first of them are meteorological data for temperatures, relative humidities, rainfall and solar radiation, possibly also concentration of acidforming gases in the atmosphere. This type of data can be obtained from meteorologists in the form of so-called TRY (Test Reference Year) data which present certain average values over a sufficiently long time period. The second type of boundary conditions involves water content in the underground soil close to the studied building. These data can be obtained again by on site analysis [3]. The second type of input parameters are hygric and thermal transport and storage parameters of the materials of the structure which appear in water and heat mass balance equations implemented in the models. In case of study of coupled moisture and heat transport, these parameters include moisture diffusivity and diffusion coefficient of water vapour, sorption isotherms and water retention curves that can be optionally expressed as moisture potential curves, thermal conductivity and specific heat capacity. Within the application of computational modelling, one must take into account that all above given parameters are functions of both temperature and moisture. Therefore, there is a need to determine these parameters in dependence on moisture and temperature changes what is highly time-consuming, and for some types of materials practically unfeasible. On that account, one must assume diversion of computational simulations from reality, and the computer codes must be calibrated and validated using experimentally measured moisture and moisture profiles. Only after this validation is done, the computational modelling can be applied with sufficient accuracy. Specific problem for validation of computational codes solving the coupled moisture and heat transport represent composite materials and structures, where interface resistances crucially affect the reliability of modelling. Most classical models of heat and moisture transport in porous materials do not deal with moisture transport across interfaces between two porous materials in an explicit way. In practical calculations, the simplest solution is commonly used. It is assumed that there is ideal contact between the two materials, which means the equality of temperatures and moistures, taken as limits on the right and the left to the interface. However, this requirement can sometimes lead to dramatic but unclear moisture profiles in the regions close to the interface, particularly if the WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


201

bordering materials have considerably different transfer coefficients. Therefore, the condition that the macroscopic capillary pressure must be continuous across the interface is adopted in more sophisticated models, implying an ideal hydraulic contact. Then, at the interface between two different materials a moisture jump across the interface appears, which can be theoretically determined on the basis of the measured water accumulation functions of both materials. In the case a non-ideal hydraulic contact is formed on the interface due to the different pore size distributions of the adjacent porous materials, a jump of capillary pressure along the interface characterized by the interface permeability can appear. Typical example of non-ideal hydraulic contact of two materials represent masonry, where the water transport properties of walling blocks and mortar are different due to the differences in their pore size distribution. In order to decide which of the particular transport models is realistic for a specific interface, an experimental analysis is required. On that account, an experimental analysis of temperature and moisture transport across the material interfaces in stone masonry is studied in this paper to reveal the interface effects on the moisture and temperature distribution.

2 Semi-scale experiments The semi-scale experiments are presently very popular in verification and calibration of HAM (Heat, Air, and Moisture transport) models since they allow monitoring of hygrothermal changes in the studied structures in more detailed way and the certain specific cases can be studied what is not very common in the case of full test house measurement [4, 5]. The semi-scale measuring system for determination of temperature and moisture fields is designed in such a way that it simulates conditions, which are as close as possible to the real conditions on building site, but it still maintains its laboratory character, so that the expenses are kept considerably lower compared to a real test house. Also the accuracy of applied measuring methods for moisture content and temperature measurement is much higher compared to insitu measurements [6].

3 Experimental Within the performed experiments, the moisture and heat transport in the fragment of argillite wall was studied. Two separate semi-scale experiments were designed and performed in order to evaluate the effect of interface resistances on the transition between the argillite walling blocks and mortar. Within the first experiment, the heat transport was simulated, whereas the relative humidity was maintained on constant level. In the second experimental arrangement, relative humidity transport was monitored at constant temperature conditions. 3.1 Studied masonry and inbuilt materials Two materials typical for historical masonry in Central European territory were used. The researched fragment of the stone masonry consisted of argillite blocks WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)

202 Materials Characterisation V and lime mortar with pozzolana admixture based on calcined kaoline mixed with milled mudstone. The composition of applied mortar is given in Table 1. Table 1: lime hydrate (kg) 200

Composition of applied plaster. pozzolana (kg) 50

sand 0 – 4 mm (kg) 750

w/d ratio (-) 0.23

In Table 1, w/d represents the water/dry substance ratio. Lime hydrate is product of lime kiln Čertovy schody, Inc., Czech Republic, whereas the silicious sand is coming from sand-pit Hlavačov, Czech Republic. The applied argillite is coming from quarry Džbán, Czech Republic. It is highly heterogeneous fine-pore material with anisotropic structure. It is characteristic by diversity in composition, whereas its main constituents are illite, kaolin, minerals of SiO2, spar and mica. Basic physical properties of both masonry materials are given in Table 2. Table 2: material mortar argillite

Basic properties of masonry materials.

bulk density (kg/m3) 1688 1353

matrix density (kg/m3) 2560 2235

total open porosity (-) 0.34 0.39

3.2 Experimental arrangement, measuring technology For simulation of moisture and heat transport in the investigated structure, a semi-scale system was used. The device (Figure 1) consists of two climatic chambers for simulation of climatic conditions, connecting tunnel, where the investigated structure is placed, and commercial devices for continuous monitoring of field variables as relative humidity, temperature, liquid moisture content, salinity, heat flux, etc. It is also equipped with optional sprinkling device for rainfall simulation. In the presented experiments, monitoring of relative humidity and temperature changes in the specific places of studied structure was done. For this purpose, sophisticated technique from Ahlborn was used. The accuracy of particular sensors was as follows: ±2% for capacitive relative humidity sensors applicable in the range of humidities 5-98%, for temperature sensors ± 0.4°C in the temperature range from –20°C to 0°C and ± 0.1°C in the range from 0°C to 70 °C. For testing the tightness of climatic chamber system, the anemometers for air flow velocity measurements were used. The whole measuring system was operated by a computer, including the climatic data entry into the particular climatic chambers. The details on the system including the measuring technology and sensors calibration can be found in [6]. The measuring process can be divided into several basic phases: walling the investigated structure, sensors’ installation into the built wall fragment, sample WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


Figure 1:

203

Climatic chamber system for semi-scale experiments.

positioning into the tunnel between the climatic chambers, climatic loading simulation, temperature and relative humidity monitoring, data evaluation. The walling of the studied structure was done in the standard way using dry mortar mix and wet technological process. The sensors for monitoring temperature and relative humidity were placed to the investigated construction to beforehand bored holes. The upper part of the bore opening was closed by silicon sealing. Placing of the sensors was done regarding to the study of interface moisture and heat transport (Figure 2). For a proper setting of climatic conditions, it was necessary to achieve nearsteady-state conditions in the studied wall. Therefore, there was necessary to dry freshly built wall to remove technological water. The drying was done at 50°C and 10% of relative humidity for one month. After that, the first experiment simulating the heat transport at constant moisture conditions was started. On one side of the studied wall, constant temperature of -9.5°C was maintained, whereas on the other side of the masonry fragment, constant temperature of 25°C was simulated. The relative humidity varied during the experiment in the range of ±10%, typically between 40 and 50% in respect to performance of climatic chambers. Duration of the first experiment was 84 days.



Figure 2:

Sensors’ positioning and wall arrangement (dimensions in mm).

The second experiment was designed to simulate gaseous moisture transport without temperature effects. Also in this case, the climatic chambers’ parameters were limiting factor. On one side of the wall, relative humidity of 45% was simulated, whereas on the opposite wall side, 95% was reached. The temperature varied within the range of 25 to 28°C. Duration of this experiment was 81 days.

4 Results and discussion The temperature profiles measured in the first semi-scale experiment are shown in Figures 3–6. The results are presented in two cross sections (A-A, B-B), according to the scheme in Figure 2. The measured temperature profiles in Figures 3–6 are more or less linear. This indicates that steady state heat transport was reached. We can also observe the expected jumps in temperatures at the contacts of measured wall with ambient environment. They are caused by thermal resistances between the material surface and ambient air. The thermal resistances between the argillite blocks and mortar are manifested as well. The effects of interface resistances are not so distinct as the surface resistances as solid-solid heat transfer can be generally realized in a more ideal way than solid-gas transfer. Nevertheless, their significance is perceptible. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


Figure 3:

Figure 4:

205

Temperature profiles in cross section A-A (first experiment).

Temperature profiles in cross section A-A (end of the first experiment).



Figure 5:

Figure 6:

Temperature profiles in cross section B-B (first experiment).

Temperature profiles in cross section B-B (end of the first experiment).

The relative profiles measured in the second semi-scale experiment are shown in Figures 7 and 8.



207

Figure 7:

Relative humidity profiles in cross section A-A (second experiment).

Figure 8:

Relative humidity profiles in cross section B-B (second experiment).


208 Materials Characterisation V Apparently, in case of simulation of water vapour transport, similar features as in heat transport simulation are observed. The experimental relative humidity profiles reveal significant effects of interface resistances on gaseous moisture transport at the contact of argillite blocks and mortar. The surface resistances remarkably affect the water vapour distribution in the studied masonry as well. The steep parts of relative humidity profiles are highly important for calibration and validation of HAM models of transport phenomena in multilayered systems. For demonstration of temperature conditions within the water vapour transport simulation, Figure 9 is introduced. Clearly, the effect of interface resistances in the stone-mortar-stone transition zone and surface resistances can be observed even in the case of low temperature differences on opposite wall sides.

transition zone

Figure 9:

Temperature distribution in cross section B-B (second experiment).

5 Conclusions Experimental investigation of water vapour and heat transport across the interfaces in stone masonry was done in the conditions of 1-D semi-scale experiment. The climatic loading of the studied structure was chosen in such a way that water vapour transport under isothermal conditions and heat transport at constant relative humidity were simulated. The data obtained in both experiments confirmed the significant effect of both interface and surface resistances on water vapour and heat propagation, thus the necessity to include these effects in computational modelling of coupled moisture and heat transport. On the basis of measured data, determination of heat and water vapour resistances between particular materials’ layers is possible using methods of WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


209

inverse analysis. Also, the surface resistances on the contact of studied building structure with ambient air can be accessed. In this way, the results can contribute to higher accuracy of future computational modelling. Neglecting the behaviour of the transition zone between the particular materials forming stone masonry structures, on the other hand, may lead to possible improper design of renovation or construction solutions and subsequent material and structural damage.

Acknowledgement This research has been supported by the Czech Ministry of Education, Youth and Sports, under project No MSM 6840770031.

References [1] Grunewald, J., Diffusiver und konvektiver Stoff- und Energietransport in kapillarporösen Baustoffen, Ph.D. Thesis, TU Dresden, Dresden, 1997. [2] Černý, R. & Rovnaníková, P., Transport Processes in Concrete, 1st ed., Spon Press: London, 547 pp., 2002. [3] Pavlík, Z., Michálek, P., Pavlíková, M., Kopecká, I., Maxová, I. & Černý, R., Water and Salt Transport and Storage Properties of Mšené Sandstone. Construction and Building Materials, 22(22), pp. 1736-1748, 2008. [4] Krus, M., Rösler, D. & Sedlbauer, K., New model for the hygrothermal calculation of condensate on the external building surface. Research in Building Physics and Building Engineering, Fazio, Ge, Rao, Desmarais (eds.), Taylor & Francis Group, London, pp. 329-333, 2006. [5] Pavlík, Z., Mihulka, J., Žumár, J. & Černý, R., Experimental monitoring of moisture transfer across interfaces in brick masonry. Structural Faults and Repair 2010 [CD-ROM], Engineering Technics Press Edinburgh, Edinburgh, 2010. [6] Pavlík, Z., Pavlík, J., Jiřičková, M. & Černý, R., System for Testing the Hygrothermal Performance of Multi-Layered Building Envelopes. Journal of Thermal Envelope & Building Science, 25(1), pp. 239-249, 2002.



Section 6 Mechanical characterisation and testing



213

Tension/compression test of auto-body steel sheets with the variation of the pre-strain and the strain rate G. H. Bae & H. Huh School of Mechanical, Aerospace and Systems Engineering, KAIST, Republic of Korea

Abstract This paper investigates the tension/compression hardening behaviour of autobody steel sheets with the variation of the pre-strain and the strain rate. To conduct tension/compression tests with the variation of the pre-strain and the strain rate, an experimental method was established by using a newly developed clamping device to suppress buckling of a specimen. The clamping device provides the supporting force from compression-type coil springs during the test with a conventional dynamic material fatigue testing machine. From experiments, the tension/compression hardening behaviour was observed with the variation of the pre-strain and the strain rate. Effects of the pre-strain and the strain rate on the hardening behaviour were also investigated based on the tension/compression test results. Keywords: tension/compression test, clamping device, pre-strain, strain rate.

1 Introduction Spring-back caused by the elastic recovery of the residual stress inside a formed part has been one of the most significant sources of defects in the sheet metal forming process in recent years. Spring-back predictability of numerical simulation, however, has not been satisfactory because its performance is not good enough to calculate accurate residual stress during the forming process. To improve spring-back predictability in numerical simulation, many researchers have been trying to use accurate information of the hardening behaviour of sheet materials in finite element analysis. The hardening behaviour of sheet materials


214 Materials Characterisation V has been measured with various mechanical testing methods in the quasi-static region [1–3]. It has been successful to predict spring-back accurately with numerical simulation by applying the precise hardening behaviour obtained from tension/compression tests at the quasi-static state [4, 5]. Numerical simulation still cannot accurately predict the final shape of the formed part after spring-back for a higher strain rate because of lack in information about the tension/compression hardening behaviour at high strain rates. To enhance spring-back predictability in numerical simulation for actual formed parts, the accurate hardening behaviour should be applied in finite element analysis by performing tension/compression tests with the variation of the strain rate as well as of the pre-strain. This paper investigates the tension/compression hardening behaviour of autobody steel sheets with the variation of the pre-strain and the strain rate. To conduct tension/compression tests, a simple clamping device was newly developed to suppress buckling of a specimen during the compression loading. The clamping device provides the side force from compression-type coil springs. The compression-type coil springs were selected for the desired clamping pressure calculated by the plate buckling theory adopted by Cao and Wang [6]. Based on the secant formula and the Euler method adopted by Boger et al. [2], specimen dimensions were also designed to prevent buckling along the width direction in the gauge section and along the longitudinal direction in the unclamped region. The strain in the gauge region of the specimen was measured directly by the digital image processing technique. The hardening behaviour of auto-body steel sheets was investigated precisely based on the tension/compression test results with the variation of the pre-strain and the strain rate.

2 Test preparation To obtain a larger compressive strain range in the tension/compression test, buckling of a specimen should be prevented by using a properly-designed specimen and by imposing a sufficient clamping force. Fig. 1 shows three representative buckling modes in the tension/compression test: (1) buckling in the thickness direction in the gauge region (T-buckling); (2) buckling in the

Figure 1:

Three representative buckling modes in the tension/compression test.



215

unclamped region (L-buckling); (3) buckling in the width direction in the gauge region (W-buckling). T-buckling mode can be suppressed by imposing a sufficient clamping force on a specimen with a special clamping device. Lbuckling and W-buckling modes can be prevented by designing an appropriate specimen shape. 2.1 Specimen shape Boger et al. [2] used the secant formula and the Euler method to calculate the attainable compressive strain for L-buckling and W-buckling modes, respectively. When the flow curve is expressed by the Swift model, the attainable compressive strain for two buckling modes can be calculated as follows: 1 é ù êæç B ö÷n ú eL = e0 êç ÷÷ -1ú êèçW ø ú ëê ûú (1) eW =

p 2 nW 2 - e0 3G 2

(2)

where B and W are the specimen width in the gripping region and the gauge region, respectively. G is the gauge length of the specimen. ε0 and n are the plastic strain for the yield stress and the hardening exponent in the Swift model, respectively. If the gauge length of the specimen and the coefficients of the Swift model are given, the attainable compressive strains are a function of the gauge width. More detailed formulation procedure can be found in the reference [2]. The specimen design was performed using the above equations formulated for the attainable compressive strain of L-buckling and W-buckling. The selected steel sheets are SPCC and DP590 which are commonly used auto-body steel sheets. Strains at uniform elongation before necking initiates are approximately 0.15 and 0.12 for SPCC and DP590, respectively. These values are assigned as the maximum pre-strain for tension/compression tests without necking of a specimen. Table 1 shows the coefficients of the Swift model and r-values for SPCC and DP590 at a quasi-static state. A specimen shape for the uniaxial tensile test [7, 8] was utilized as reference specimen dimensions for tension/compression tests as shown in fig. 3(a). In order to reduce the required clamping force to suppress T-buckling, the gauge length is first reduced from 30 mm to 20 mm. The small clamping force is favourable for the reliable tension/compression test, which can reduce the frictional and biaxial effect on a specimen. The attainable compressive strain was plotted with respect to the gauge width as shown in fig. 2. From the plotted curves, the optimal gauge width can be selected as 8.8 mm for SPCC and 8.2 mm for DP590, respectively. To ensure a stable testing region for the two steel sheets, the final gauge width can be determined to be 8.8 mm as shown in fig. 3(b).


216 Materials Characterisation V Table 1:

Coefficients of the Swift model and r-values of SPCC and DP590. K [MPa]

ε0

n

r0

r45

r90

R

SPCC

555.772

0.01121

0.253

1.381

1.094

1.697

1.317

DP590

1079.261

0.00879

0.220

0.704

0.784

0.948

0.805

Attainable compressive strain

Material

0.4

L-buckling SPCC DP590

0.3

W-buckling SPCC DP590

0.2

SPCC, 0.15 DP590, 0.12

0.1 0.0

0

5

10

15

20

25

Width of the gauge region [mm]

Figure 2:

Attainable compressive strain with respect to the width at the gauge region for SPCC and DP590.

(a)

(b) Figure 3:

Specimen dimensions for the uniaxial tensile test and the tension/compression test: (a) uniaxial tension test [7, 8]; (b) tension/compression test (suggested).

2.2 Clamping force Cao and Wang [6] proposed an equation to calculate the blank holding force to suppress flange wrinkling in the sheet metal forming process. To calculate the clamping force to suppress T-buckling of a specimen, the plate buckling theory was also employed by assuming the gauge region of a specimen to be a rectangular plate model. The critical normal force required to suppress buckling of a specimen can be calculated by WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)

Materials Characterisation V n +1 éï ù æ 2u x ö÷ïüï êì ú ï ç e ln 1 ÷ í ý ç ê 0 ú çè G ÷øïþ 3KGWt êïîï ú ï F= ê ú n +1 n 4(n + 1)d ê 2 æ ct ú ö÷ æ 1 æ ö ö 1 t t 1 ç ÷ ç ÷ ê ú ç ê- G èçç 2 + e0 èçç m 2 d + 2 ø÷÷ø÷÷÷ èçç m 2 d + 2 ø÷÷ tan (md )ú ë û

217

(3)

2u x ( l - u x ) 2p ,d= G 2 u p 1+ 2R x where t and ux are the thickness of a specimen and a certain edge displacement. K, n and ε0 are coefficients of the Swift model. For the tension/compression test, the specimen thickness and width change in the gauge region during deformation. These values can be calculated based on anisotropy of material. where c =

1+ R

, m=

W = W0 -

t = t0 -

æ 2u ö W0 ln çç1- x ÷÷÷ 1 + R çè G ø

(4)

t0 R æç 2ux ö÷ ln ç1÷ 1 + R çè G ø÷

(5)

The critical normal force defined in eqn. (3) is now used to calculate the critical clamping force to suppress T-buckling in the tension/compression test. More detailed formulation procedure of the critical normal force can be found in the reference [6]. Fig. 4 shows the critical clamping force with respect to the compressive plastic strain. From the plotted curves, the required clamping force for SPCC and DP590 can be determined to be 2.003 kN and 2.275 kN, respectively.

Clamping force [kN]

10

G20W8.8 SPCC 8 DP590

DP590 0.12

SPCC 0.15

6 4 2 0 0.00

0.05

0.10

0.15

0.20

Compressive strain

Figure 4:

Required clamping force with respect to the compressive strain for SPCC and DP590.

2.3 Design of a new clamping device Based on specimen dimensions and the required clamping force, a new clamping device was developed to suppress T-buckling of specimens. Fig. 5(a) gives a schematic diagram for the clamping device developed. A controllable clamping WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)

218 Materials Characterisation V system was suggested by using compression-type coil springs. Four coil springs on each side compress a specimen by fastening the bolts up to the desired deflection distance which is determined by the gap controlling bar. Before the tension/compression test starts, the gap controlling bar is uninstalled from the clamping device in order to allow thickness change of the specimen during the tension/compression test. For lubrication between the specimen and the clamping plate, a Teflon film (3M 5490 model) with a thickness of 0.09 mm is attached on the clamping plate. A compression-type coil spring was selected based on the clamping force calculated in Section 2.2 for SPCC and DP590. The maximum clamping force was selected at 4 kN to ensure performance of the device although the required clamping forces were 2.003 kN for SPCC and 2.275 kN for DP590. To provide the required clamping forces, Φ16L40 coil spring was selected for the clamping device. The maximum load and the spring constant of the selected coil spring are 1.265 kN and 0.158 kN/mm, respectively. The clamping device was manufactured with the specified dimensions based on the specimen shape and the spring size. As shown in fig. 5(b), the springloaded clamping device has a small and simple structure with light weight. The clamping force is controllable by determining the deflection distance of the coil springs. These advantages are favourable for easy and convenient tension/compression tests at high strain rates. Specimen Teflon film

Coil springs Gap controlling bar

Bolt

Nut

Clamping plate

Compression plate

(a) Figure 5:

(b)

Spring-loaded clamping device for the tension/compression test: (a) schematic diagram; (b) manufactured clamping device.

3 Experiments 3.1 Test material The tension/compression test of auto-body steel sheets was conducted with the variation of the pre-strain and the strain rate. Test materials were SPCC and DP590 with a thickness of 1.2 mm. The chemical composition of the sheet metals is presented in table 2. Table 2: Material SPCC DP590

Chemical composition of SPCC and DP590 [wt.%]. C 0.085 0.326

Mn 0.420 1.700

Si 0.040 0.123


P 0.0012 0.0162

S 0.0150 0.0009


219

3.2 Testing machine A clamped specimen is gripped by the dynamic material fatigue testing machine, INSTRON8801, to perform tension/compression tests with the variation of the pre-strain and the strain rate. The dynamic material fatigue testing machine shown in fig. 6 has a maximum stroke of ±75 mm in the vertical direction actuated by a hydraulic system. The load cell installed in the testing machine can measure a load of up to ±100 kN. The gripper is also actuated by a hydraulic system.

Figure 6:

Tension/compression testing system utilizing the dynamic material fatigue testing machine, INSTRON8801, and the spring-loaded clamping device.

True strain in the T/C test

0.15

Actuator displacement: 3 mm Strain rate: 0.001 /sec

0.10

0.05 SPCC DP590 0.00

0

1

2

3

4

5

6

Total actuator movement [mm]

(a) Figure 7:

(b)

Strain measuring method using the digital image processing technique with the high speed camera: (a) recording of the deformation history; (b) measured strain during the tension/compression test.

3.3 Strain measurement Because the specimen is fully covered by the clamping plates, conventional extensometers cannot be employed to measure the strain in the side of the gauge region during the tension/compression test. As an alternative, a digital image processing technique was utilized to measure the strain directly. Before the tension/compression test, a speckled pattern was sprayed on the side wall of the specimen as shown in fig. 7(a). A high speed camera captures a series of frames for deformation during the tension/compression test. The strain can be measured WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)

220 Materials Characterisation V from a graph in fig. 7(b) obtained by analyzing recorded frames using the digital image processing technique. 3.4 Tension/compression test Stable tension/compression testing conditions were established by comparing the response of the testing machine with the variation of the actuator displacement and the strain rate. Table 3 presents experimental conditions for the tension /compression test. From the experiments, load-displacement curves and strains in the gauge region were determined for various pre-strains and strain rates. The tension/compression hardening curves can be obtained after correction procedures to subtract the frictional and biaxial effects induced by the clamping force. Table 3:

Material

SPCC

DP590

Experimental table for tension/compression tests of SPCC and DP590. Displacement [mm] 1.0 2.0 3.0 4.0 1.0 2.0 3.0

0.001 O O O O O O O

Strain rate [/sec] 0.01 0.1 O O O O O O O O O O O O O O

1.0 X X O O X X O

3.5 Correction of acquired load All acquired stress−strain results require corrections for frictional and biaxial effects induced by the clamping force, Fc , because the specimen is compressed by the spring-loaded clamping device in the thickness direction to suppress T-buckling. The measured force from the load cell, Fm , is the addition of the force from the specimen deformation, Fd , and the frictional force, Ff .

Fd = Fm - Ff

(6)

The friction behaviour is represented by the Coulomb friction law as follows [1, 2]:

Ff = m Fc

(7)

The frictional conditions can be changed according to the material, the geometrical change of a specimen and the accumulated damage to the Teflon film. The friction coefficient is generally known to be in the range of 0.03~0.09 when the Teflon film is used for tension/compression tests [2–4]. The friction coefficients for SPCC and DP590 are selected as 0.06 and 0.08, respectively, to correct the acquired curve obtained from the tensile test with clamped into a WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


221

corrected curve by comparison with the acquired curve obtained from a tensile test with unclamped. In order to subtract the through-thickness stress, σt , induced by the clamping force, the von Mises yield function is utilized to calculate the effective stress corrected from the biaxial effect.

1é 2 (sm - st ) + sm2 + st2 ùúû ê ë 2

s=

(8)

where σm is the measured stress from the load cell. After both frictional and biaxial corrections, the flow curve of the clamped specimen agrees well with that of the unclamped specimen as shown in fig. 8.

400

SPCC

1000

Strain rate = 0.001 /sec Clamping force = 2.003 kN  = 0.06

300 200 Acquired curve (unclamped) Acquired curve (clamped) Corrected curve

100 0 0.00

0.05

0.10

True stress [MPa]

True stress [MPa]

500

800

DP590 Strain rate = 0.001 /sec Clamping force = 2.275 kN  = 0.08

600 400 Acquired curve (unclamped) Acquired curve (clamped) Corrected curve

200 0 0.00

0.15

0.05

True strain

(a) Figure 8:

Frictional and biaxial corrections in the uniaxial tension state: (a) SPCC, 0.001/sec; (b) DP590, 0.001/sec. 500

SPCC, 0.001 /sec

250

Simple tension Displacement (pre-strain) 1 mm (0.040) 2 mm (0.080) 3 mm (0.120) 4 mm (0.162)

0 -250 -500 0.00

0.05

0.10

0.15

SPCC, 0.01 /sec

250


0 -250 -500 0.00

0.20

0.05

True strain


0 -250

0.10

0.15

True strain

0.20

SPCC, 1.0 /sec

250

Simple tension Displacement (pre-strain) 3 mm (0.116) 4 mm (0.159)

0 -250 -500 0.00

0.05

0.10

0.15

0.20

True strain

(c) Figure 9:

0.20

True stress [MPa]

True stress [MPa]

SPCC, 0.1 /sec

0.05

0.15

(b) 500

250

-500 0.00

0.10 True strain

(a) 500

0.15

(b)

True stress [MPa]

True stress [MPa]

500

0.10

True strain

(d)

Tension/compression test results of SPCC at various strain rates: (a) 0.001/sec; (b) 0.01/sec; (c) 0.1/sec; (d) 1.0/sec.


222 Materials Characterisation V 1000

DP590, 0.001 /sec

500

Simple tension Displacement (pre-strain) 1 mm (0.032) 2 mm (0.070) 3 mm (0.112)

0 -500 -1000 0.00

0.05

0.10

True stress [MPa]

True stress [MPa]

1000

DP590, 0.01 /sec

500


0 -500 -1000 0.00

0.15

0.05

True strain

(a) DP590, 0.1 /sec

500


0 -500 -1000 0.00

0.05

0.10

DP590, 1.0 /sec

500

Simple tension Displacement (pre-strain) 3 mm (0.106)

0 -500 -1000 0.00

0.15

0.05

True strain

Tension/compression test results of DP590 at various strain rates: (a) 0.001/sec; (b) 0.01/sec; (c) 0.1/sec; (d) 1.0/sec. 500

SPCC, 1 mm

250

Strain rate [/sec] 0.1 0.01 0.001

0 -250 -500 0.00

0.05

0.10

0.15

SPCC, 2 mm

250

Strain rate [/sec] 0.1 0.01 0.001

0 -250 -500 0.00

0.20

0.05

True strain

Strain rate [/sec] 1.0 0.1 0.01 0.001

0 -250

0.10

0.15

0.20

True stress [MPa]

True stress [MPa]

250

0.05

0.20

SPCC, 4 mm

250

Strain rate [/sec] 1.0 0.1 0.01 0.001

0 -250 -500 0.00

0.05

True strain

(c) Figure 11:

0.15

(b) 500

SPCC, 3 mm

-500 0.00

0.10 True strain

(a) 500

0.15

(d)

True stress [MPa]

True stress [MPa]

500

0.10

True strain

(c) Figure 10:

0.15

(b) 1000 True stress [MPa]

True stress [MPa]

1000

0.10

True strain

0.10

0.15

0.20

True strain

(d)

Tension/compression test results of SPCC at various actuator displacements (pre-strains): (a) 0.001/sec; (b) 0.01/sec; (c) 0.1/sec; (d) 1.0/sec.



1000

DP590, 1 mm

500

Strain rate [/sec] 0.1 0.01 0.001

0 -500 -1000 0.00

0.05

0.10

True stress [MPa]

True stress [MPa]

1000

DP590, 2 mm

500

Strain rate [/sec] 0.1 0.01 0.001

0 -500 -1000 0.00

0.15

0.05

True strain

0.10

0.15

True strain

(a)

(b)

1000 True stress [MPa]

223

DP590, 3 mm

500

Strain rate [/sec] 1.0 0.1 0.01 0.001

0 -500 -1000 0.00

0.05

0.10

0.15

True strain

(c) Figure 12:

Tension/compression test results of DP590 at various actuator displacements (pre-strains): (a) 0.001/sec; (b) 0.01/sec; (c) 0.1/sec.

4 Hardening behaviour True stress−true strain curves were obtained after frictional and biaxial corrections, as shown in figs. 9 and 10. Hardening curves expand with the increase of the pre-strain at various strain rates. To investigate the strain rate effect on the hardening behaviour, true stress−true strain curves were redrawn with the variation of the strain rate as shown in figs. 11 and 12. Hardening curves also expand with the increase of the strain rate at various pre-strains. It is well known that the strain and the strain-rate hardening are caused by micro-structure changes inside the materials. To confirm this general expectation, Huh et al. [9] demonstrated by observing TEM experimental results that the mechanism of the strain-rate hardening is the change of the dislocation structures and the increase of the dislocation density. The mechanism of the hardening behaviour during tension/compression loading also can be explained by the micro-structure change of steel sheets caused by the pre-strain and the strain rate effect. From the experiments, it is clear that the tension/compression hardening behaviour is changed by the strain rate as well as by the pre-strain. This means that the prestrain and the strain rate effect on the tension/compression hardening behaviour should be considered simultaneously to improve spring-back predictability in numerical sheet metal forming simulation.



5 Conclusions This paper investigates tension/compression hardening behaviour by performing tension/compression tests using a newly developed spring-loaded clamping device. Contributions in this paper are summarized as follows: 1) A specimen shape was designed with the secant formula and the Euler method adopted by Boger et al. [2] to prevent L-buckling and W-buckling in the desired tension/compression test range. The clamping force to suppress Tbuckling was calculated with the plate buckling theory. A spring-loaded clamping device was newly developed to suppress T-buckling in the tension/compression test. From experiments, it is noted that the dimension of a specimen and the choice of compression-type coil springs are appropriate to execute reliable tension/compression tests. 2) The strain in the gauge region of a clamped specimen was measured directly by using a digital image processing technique in order to overcome the inherent structural problems in the use of conventional contact-type extensometers. The measuring scheme using the digital image processing technique can provide accurate strain history in the gauge region of a clamped specimen during tension/compression tests. 3) With the test conditions established, hardening curves were reliably obtained for various pre-strains and strain rates. The tension/compression hardening curves were expanded with the increase of the strain rate as well as with the increase of the pre-strain. The mechanism of the pre-strain and strain rate hardening was explained by micro-structure changes, such as dislocation structures and dislocation density. Experiments provide good information of the hardening behaviour with respect to the pre-strain and the strain rate. 4) The newly developed clamping device can be directly applied to tension/compression tests at high strain rates because of its compactness and convenience to control the clamping force with various tensile testing machines. This device guarantees the response of the clamping pressure in real time when the clamping condition is changed by the specimen deformation during tension/compression tests.

References [1] Lee, M. G., Kim, D., Kim, C., Wenner, M. L., Wagoner, R. H. & Chung, K., Spring–back evaluation of automotive sheets based on isotropic–kinematic hardening laws and non–quadratic anisotropic yield functions, part II: characterization of material properties. Int. J. Plast., 21, pp. 883–914, 2005. [2] Boger, R. K., Wagoner, R. H., Barlat, F., Lee, M.–G. & Chung, K., Continuous, large strain, tension–compression testing of sheet material. Int. J. Plast., 21, pp. 2319–2343, 2005. [3] Cao, J., Lee, W., Cheng, H. S., Seniw, M., Wang, H. P. & Chung, K., Experimental and numerical investigation of combined isotropic–kinematic hardening behavior of sheet metals. Int. J. Plast., 25, pp. 942–972, 2009.



225

[4] Lee, M. G., Kim, D., Kim, C., Wenner, M. L. & Chung, K., Spring–back evaluation of automotive sheets based on isotropic–kinematic hardening laws and non–quadratic anisotropic yield functions. part III: applications. Int. J. Plast., 21, pp. 915–953, 2005. [5] Taherizadeh, A., Ghaei, A., Green, D. E. & Altenhof, W. J., Finite element simulation of springback for a channel draw process with drawbead using different hardening models. Int. J. Mech. Sci., 51, pp. 314–325, 2009. [6] Cao, J. & Wang, X., An analytical model for plate wrinkling under tri–axial loading and its application. Int. J. Mech. Sci., 42, pp. 617–633, 2000. [7] Huh, H., Kim, S. B., Song, J. H. & Lim, J. H., Dynamic tensile characteristics of TRIP–type and DP–type steel sheets for an auto–body. Int. J. Mech. Sci., 50, pp. 918–931, 2008. [8] Huh, H., Lim, J. H. & Park, S. H., High speed tensile test of steel sheets for the stress−strain curve at the intermediate strain rate, Int. J. Automotive Technol., 10(2), pp. 195–204, 2009. [9] Huh, H., Yoon, J. H., Park, C. G., Kang, J. S., Huh, M. Y. & Kang, H. G, Correlation of microscopic structures to the strain rate hardening of SPCC steel. Int. J. Mech. Sci., 52, pp. 745–753, 2010.




227

Definition of averaged elastic-plastic characteristics of sandwich panel structures I. I. Zakirov, V. N. Paimushin & I. M. Zakirov A. N. Tupolev State Technical University, Kazan, Russia

Abstract Analytic formulas for averaged elastic and strength characteristics of folded structure’s determination are identified. They are based on the introduction of a hypothesis about momentless action of its elements (web covers) in a charge operation and on phased deletion from action ex-post buckling. Dimensionless coefficients contained in the structural formulas, which appear on the solution process of formulated problems, are to be determined from experimental evidence on tension, compression and pure shear in two-planes. In a tensioncompression process the folded structures should be considered heterogeneous material with averaged plasto-elastic behaviors, which has relations between averaged voltages and averaged appropriate deformations and has a deformations stepwise change area (conditional plastic flow). Keywords: folded structures, cell of cyclicity, averaged elastic and strength characteristics, modulus of elasticity, deformation curve, structural formulas, theoretical and experimental method.

1

Introduction

In many cases, the use of composite materials is the most efficient in constructions (three-layer or multilayer) with cores, which have good mechanical strength, rigidity, vibration, sound and heat insulation characteristics. These characteristics of cores of some structures have been well studied today. A large number of publications, a detailed analysis of which is contained, in particular, in a review article [1] are devoted to their research. Most often in three-layer and multilayer structures, the cores have a honeycomb ([2, 10] etc.) structure. In the calculations of such structures the real cores are usually replaced by some homogeneous continuum, which averaged WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line) doi:10.2495/MC110201

228 Materials Characterisation V mechanical and thermal characteristics are determined by the principle of equivalence of real and replacing core [3–5]. At such replacement of real honeycomb with orthotropic continuum for the strength calculations averaged elastic module in the tangential directions and the shear modulus in the tangential plane are set equal to zero, and the elastic modulus in the transverse direction and shear module in the transverse plane is determined by theoretical or experimental methods [8, 10]. Formulas for their determination are given, in particular, in references [3, 4] that, as a first approximation of formulas, have been obtained by the energy method [3], or the method of displacements and forces [6]. Along with these, there are other theoretical methods for determining the averaged elastic characteristics of periodic structures based, in particular, on the method of averaging the solution of problems by Bakhvalov and Ponasenko [7]. In recent years, ways of production of the folded structures from flat sheets cores [11, 12] develop very intensively. Such cores compared with honeycomb cores have a number of fundamental differences and advantages. In this paper we consider one of these folded structures. We develop relevant theoretical and experimental method to determine the homogenized elastic and strength characteristics of which. 1.1 Formulation of the problem Consider folded structure with a quadriradiare core [11], formed from a flat sheet blank with thickness t by means of its gradual transformation into a state of

relief, as shown in Figure 1a. We assume that, in general, l1  l2 , 1  2 and all four sloping walls 1,1, 2, 2 of the periodicity cells isolated from the core are a parallelogram, symmetrically placed relative to the plane the middle planes

x01 z . Therefore,

  k  of each of them reasonable classified as a local oblique

xk 0k yk and xk 0k yk , in which (Fig. 1a) xk  xk and the radius vectors of the points   k  given by a representation coordinate system

r    xk e1   yk e2  , r  k

k

In them unit vectors

k

k

k k  xk e1   yk e2 

(1.1)

ei k  , ei k  satisfies the equations 

ei  ei   1, ei  ei k

k 

k

k 

 1, e1  e2   e1  e2   cos  k k

k

k

k

(1.2)

which according to Fig. 1a and 1b with unit vectors i , j , k of orthogonal Cartesian coordinate system related with dependencies



229

ie1 k   ie1 k   cos  k , ie2 k   ie2 k   sin 



je2 k   cos , je2 k    cos , je1 k   je1 k   0 



ke11  ke11   sin 1 , ke1 2  ke1 2    sin  2 



(1.3)

ke2 k   ke2 k   0 

(а)

(b) Figure 1. Parameterizations (1.1) correspond to the basis vectors (for the elements

1 and 2 compiled following relations must be marked with strokes) r1   r  k

k

xk  e1  , r2   r  k

k

k

yk  e2


k

(1.4)

230 Materials Characterisation V co- and contravariant components of metric tensor   aii   ri  ri   1, a12   a21  r1  r2   cos  k k

k

aiik  

k

k

k

k

k 

k

(1.5)

k  12

aii cos  1 a , a12k   a21k      2 k 2 a k  sin k a k  sin  k

k  k 

k  k 

where a k   a11 a22  a12 a12  1  cos

2

(1.6)

k  sin 2 k .

For the displacement vectors of core walls we accept following representations U  k   u  k  e1k   v k  e2k   w k  m  k   z k    k  e1k     k  e2k 





 t 2  z k   t 2

(1.7)

eik  are the reciprocal basis vectors satisfying the conditions

where

eik  e j    ij (  ii  1,  ij  0 where i  j ), aijk   eik  ejk  , m  k  are the unit k

vectors

normal

to

the

plane

 k  ,

satisfying

the

equations

ei  m    0, m   m    1 . k

k

k

k

According to the Kirchhoff-Love model, with an average bending for the functions 

k 

and

 k 

will take place following dependences.

  k   w,xk  ,   k   w,yk  where

 , x    

xk . By using these relations for the covariant

components of the tangential deformations are established following relations.

 ij    ij k   z k  æijk  zk

(1.8)

where

1 2 k k k k   u, y  v,x   w,x  w,y 

1 2

11 k   u,xk   w,xk  w,xk  ,  22 k   v,yk   w,yk  w,yk  k 

212

k  k  æ11  w,xxk  , æ22k   w,yyk  , æ12  w,xyk 

(1.9)

(1.10)

Under the assumption of linear elastic behavior of core material when it is loading the contravariant components of the tension tensors in the walls with deformation tensor components

 ij  zk

 ijk 

are related by the elasticity

correlation of the following form WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


 ijk  

E z Eijml  k  k  ml 2 1 v

231

(1.11)

where E , v are the Young’s modulus of elasticity and Poisson's ratio of wall ijml

material, and E k  is the tetravalent tensor of the elastic characteristics, components of which for an isotropic material in plane stress state are defined by correlations

Eijml  vaijk  amlk   k

1  v im jl a k  a k   ailk  ajm k 2





(1.12)

Substitution of formulas (1.6) to the correlations (1.12) leads to the formulas v sin 2 k  cos 2  k 1 E1111  E2222  , E1122  E2211  , k k k k 2 sin k sin 4 k

E1112  E1121  E1211  E2111  E1222  E2122  E2212  k k k k k k k  E2221  k

cos  k , sin 4  k

E1212  E2121  E1221  E2112  k k k k

1  cos 2  k  v sin 2  k 2sin 4  k

(1.13)

We assume that under the loading of the core in the sandwich and multilayer structural elements in its walls because of their small thickness realized the flat and membranaceous stress-strain state. In this case, accumulated in them potential energy of deformation is permissible to calculate by the following formulas l d 1k k  (1.14) sin k dxk dyk Ï  k     t  11k 11 k   2 12k 12 k    22k  22 200





where d k 

ak dxk dyk  sin  k dxk dyk is the area of infinitely small element of the middle plane of k -th core wall, and the components of the deformation

 ij k  are defined by correlations (1.9). These correlations are written

in linear approximation by dropping the nonlinear terms. After the transition from a flat membrane state in the perturbed moment state possible due to a loss of stability of the flat form of equilibrium. To study the perturbed neutral equilibrium state we will use the variational equation of the Ritz method, compiled on the basis of relations (1.8)–(1.11). In the space V of a selected cell of periodicity, referred to orthogonal Cartesian coordinates x, y and z , averaged elastic properties of the core should be considered orthotropic. The elastic modulus of the first kind Ez in the direction z and the modules of the transverse shifts Gxz , G yz in the planes xz and

yz must be determined within the bounds of the model of a transversely WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)

232 Materials Characterisation V soft core. If the displacement vector of any point M  V provide an expansion U  ui  vj  wk and in cross-sections x  const , y  const , z  const introduce the averaged voltage

 xz ,  yz ,  z

the accumulated in the volume

strain energy will be determined by the expression

Ï 

1  xz xz   yz yz   z z dV 2  V

where for calculation the transverse shear strain

 xz ,  yz

(1.15)

z

and axial strain

of

the homogenized medium in the linear approximation we have the kinematic relations

 xz 

u w v w w ,  yz   , z   z x z y z

(1.16)

Using of the principle of equivalence of real and replacing conditional core with average elastic and strength characteristics requires to preparation of the equality.

Ï Ï

1

Ï

 2

Ï

1

Ï

 2

(1.17)

which will be used for further calculations.

2 Critical values of the tensions formed in the walls of the core when it is loading Assume that in walls of the core in terms of loading formed homogeneous in their middle plane tensions

 11k  ,  22k 

and

 12k  .

To produce the approximate structural formulas for using of theoretical and experimental method [8], the edges of the walls xk  0, xk  lk , yk  0, yk  d will be assumed simply supported. Whether for these or other types of loading in the walls of the core forms only compressive tensions

 11k 

and

 22k  ,

and

 12k   0 ,

when taken under

considerationт pinning the edges of the walls to construct the solution of the stability function w

k 

can be represented as

k  w k   Amn sin

m xk m yk sin lk d


(2.1)


233

where m  1, 2,..., n  1, 2,... are the wave numbers. Then with the introduction of the dimensionless parameters

lk , d

k 

 11k  it is possible to receive the formula

For definition of critical value



k 

 Kx

11 k 

 k   22k   11k  .

Dk  2E  t  k    K x Dk , Dk    sin 4  k 12 1  v 2   lk 

2

(2.2)

where k 

Kx

m 

2

  k2 n 2   4m 2 n 2  k2 cos 2  k 2

m  k  n 2

2 k

2

, Dk 

to construct the solution of the stability function w

represented as

w k   W  k  sin In a case

(2.3)

 11k   0 ,  12k   0

In general, the subcritical tension state, when

 22k   0

Dk sin 4 k

 11k    22k   0 ,

critical tension

 12k    êð 

 êð k    5.34  4 

 yk lk

sin

k 

and

can be

  yk  kk xk 

(2.4)

d

using the (2.4) for determination the minimum

it possible to receive an approximate formula

lk2   2  Dk 1  cos  k d2 

 2 1  cos  k 1  cos 2 k 

  

(2.5)

3 Determination of elastic and strength characteristics Displacement vector in a space V of core we can represent in the form

z (3.1)  u0 i  v0 j  w0 k  , 0  z  h h where u0  const , v0  const and w0  const – moving the points of the z h, with a net in-plane shear boundary plane

U

x01 z  u0  0, v0  0, w0  0  , y 01 z  u0  0, v0  0, w0  0 

and

tension-

 k  1, 2,1, 2

(3.2)

compression in the direction of the axis 01 z . We denote

U 0 k   u0 k e1k   v0 k  e2k   w0 k  m  k 


234 Materials Characterisation V displacement vectors of faces x1  x1  l1 and x2  x2  0 walls of the filler, and for the displacements U

k 

of the points M k    k  we take the

approximations

x1 1 1 x1 1  2  x2   2  U 0 , U  U 0 , U  1   U 0 l1 l1  l2   x  2  1  2  U 0   l2 

1 U  

U

 2 

(3.3)

The condition of equivalence of the real and the conventional core requires the satisfaction of the equality k k k k k (3.4) u0 i  v0 j  w0 k  U 0   u0  e1k  v0  e 2k  w0  m    k  1, 2,1, 2   

 

k 

Scalar multiplication of both sides by vectors ei

and using (1.2) (1.3), we

arrive to the dependencies

u0   u0 cos 1  w0 sin 1 , v0   u0 sin  v0 cos 1

1

u01   u0 cos 1  w0 sin 1 , v01   u0 sin  v0 cos 



u0   u0 cos  2  w0 sin  2 , v0   u0 sin  v0 cos 2

(3.5)

2

u0   u0 cos  2  w0 sin  2 , v0   u0 sin  v0 cos 2

2

In component form the vector equalities (3.4) lead to the approximating functions

u 1 

x1 1 1 x1 1 1 x1 1 1 x1 1 u0 , v  v0 , u  u0 , v  v0 l1 l1 l1 l1

 x  2 2  x  2 2 u    1  2  u0  , v   1  2  v0   l2   l2 

(3.6)

 x    x     u  2   1  2  u0 2  , v  2   1  2  v0 2   l2   l2  By the substitution into the linear kinematic relations

11 k   u,xk  ,  22 k   v,yk  , 212 k   u,yk   v,xk 

(3.7)

compute the components of deformation in the walls of the core in the membrane approximation. WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


235

Compiled on the basis of the relations, we consider the case of the periodicity cell of core by tension and compression along the axis z  w0  0, uo  v0  0  in the case of loading we can get formulas:

11 k   11 k   sin 2  k 

w0  k  k k  k  ,  22   22  12   12   0 h

E E sin 2  k w0 1111  k    E 11  1  v2 k  1  v 2 sin 4 k h E E2211  22k   11 k    v sin 2 k  cos 2 k   11k  k 2 1 v 2 E 1211  k  E 11   cos k 11k   12k   1  v2 k 

(3.8)

11 k 

(3.9)

From which follows that in the condition of tensile fracture of core ( w0  0 ) is possible because of the destruction of the adhesive layer connecting the core with a bearing layer, and while compression ( w0  0 ) due to loss of stability of the walls in her bilateral compression and shear. When we approximating the displacements of the averaged core (3.1), it formed the strains and stresses

w0 u v w ,  xz  0 ,  yz  0 ,  33  Ez 0 h h h h (3.10) u0 23 v0 13   Gxz ,   G yz h h where the homogenized elastic modules Ez , Gxz and G yz must be determined.

z 

By using the relations (3.10) in accordance with expression (1.15) for calculation Ï we arrive at the formulas in the case u0  v0  0

Ï  hd  ctg1  ctg 2  sin 33 w0

(3.11)

Ï  d  ctg1  ctg 2  sin Ez w

(3.12)

2 0

and by using relations (3.8) and (3.9), in accordance with (1.14) for calculation Ï  k  we arrive to the formulas

Ï

Ï

sin  k sin  k w0  k  1, 2,1, 2  2

(3.13)

sin 3  k w02 E td  k  1, 2,1, 2 1  v 2 2sin 3  k h

(3.14)

k 

  11k td

k 




236 Materials Characterisation V By substituting relations (3.12) , (3.14) from equation (1.17) we can obtain a formula to determine Ez

Ez 

2 sin 3  k E t  1  v 2 h  ctg1  ctg 2  k 1 sin 3 k

(3.15)

and using (3.13) from (1.17) we can obtain the following dependence

 33 

2 t  11k  sin  k sin k  h sin  ctg1  ctg 2  k 1

In the case of the maximum tensile stresses

 11k  ,  22k  ,  12k 

(3.16)

are largest in the

 k  sin 2  k sin 4 k .The process of  destruction in it begins when certain tensile strength  is reached. As in real cores, angles  k and k do not have significant difference, it is possible to 11 11 assume that when the strength reaches  1 or   2  the simultaneous failure of all

wall, in which the maximum value is

four walls of the core begins. Under these assumptions to determine the limit

 33

at which the core is destroyed we can create equality

 33 

2 t  sin  k sink  h sin  ctg1  ctg 2  k 1

(3.17)

33 From it by the experimentally determined value   determined the quantity

. When compressing of the core begins its destruction due to the destruction of the buckling of the wall, which formed the voltage

k  v sin 2 k  cos 2 k

 11k 

and cos  k

at a given dimensions

   12k   11k 

reaches a critical value



11  k 

k 

 Kx

2

t   2 E sin 2  k  t  k     Kx   (3.18) 12 1  v 2  sin 4 k  lk  12 1  v 2  sin 4 k  h 

 2E

2

Assuming further that for the real core loss of stability of a wall immediately causes a loss of stability of a second wall, we form the equation

 11k    11k  ,

from which to determine the lateral deformation at the moment of loss of stability [9] by using relations (3.9) and (3.18) follows the formula 

2  t   w0   K x     12  h   h 

2


(3.19)


1

 2

where K x  K x

237

 K x . However

 2E t   K x   2 12 1  v  sin  ctg1  ctg 2   h  33 

For given parameters E , , t h ,  k and determined value of critical stress

k of

 33 of

3

2

sin 3  k

 sin k 1

3

k

(3.20)

the core the experimentally

the derived formula (3.20) is

determined by a dimensionless coefficient K x . Elastic and strength characteristics of the core in a shift in planes XoZ YoZ are defined by analogy with the stated above.

4

Conclusions

Theoretical and the experimental methods presented in this article allow to define the averaged elastic and strength characteristics of sandwich folded fillers on the basis of use of analytical expressions for definition of the critical tensions formed in walls of a filler while loading.

References [1]

[2] [3] [4] [5]

[6] [7] [8]

[9]

Noor A. K., Burton W.S., Bert Ch. W., Computational models for sandwich panels and shells. Applied Mechanics Reviews, 1996, V. 49, ¹ 3, p. 155-199. Bersudsky V.E., Krisin V.N., Forest S., Manufacturing technology of cellular aircraft structures. - Moscow: Mashinostroenie, 1975. 216s. Aleksandrov A.Y., Bryukker L.E., Kurshin L.M., The calculation of sandwich panels / M. Oborongiz, 1960. 272s. Panin V.F., Gladkov Yu.A., Constructions with cores. Directory. Moscow: Mashinostroenie. 1991. 272s. Kryutchenko V.E., Analysis of optimum insulation properties of sandwich plates with honeycomb core. /Mechanics of Composite Materials. - 1993, T.29.- № 6.- S.835-839. Relsey S., Gellatly H. and Clark B., The shear modulus of foil honeycomb cores. Aircraft Engineering. 1958. V.30. № 356. P.294-302. Bakhvalov N.S., Ponasenko G.P., Averaging of the processes in periodic media.- Moscow: Nauka, 1984. 352 c. Sachenkov A.V., Theoretical and experimental method for studying the stability of plates and shells. Investigations on the theory of plates and shells. Kazan: Izdatel'stvo Kazan. State. Univ. - 1970, Vol. 6.7.- S. 391433. Volmir AS Stability of elastic systems. M.: Gos. out of Sci. literature. 1963. 880 sec.


238 Materials Characterisation V [10]

[11]

[12]

Akishev N.I., Zakirov I.I., Paimushin V.N., Shishov M.A., Theoreticalexperimental method for determining the averaged elastic and strength characteristics of honeycomb of sandwich structures. Mechanics of Composite Materials. - 2011 (in press). Zakirov I.M., Alekseev, K.A., Akishev N.I. Kayumov R.A., NikitinA.V., Zakirov I.I., Manufacturing of sandwich panels with folded core of polymer paper Kazan: Publishing house «Fan» 2009 Zakirov I.M., Alekseev K.A., Determining of the parameters of the four radial helical folded structure / IVUZ “Aviatcionnaya technika” 2005



239

Hot deformation and mechanical properties of P/M Al special M. Tercelj, P. Cvahte, I. Perus & G. Kugler Faculty of Natural Science and Engineering, University of Ljubljana, Slovenia

Abstract Almost pure aluminum composite, P/M Al-special, with an average grain diameter of 1μm with oxides on the grain surface was produced by a powder metallurgy route. Billets were extruded on an indirect press at various temperatures and ram-speeds. Specimens made from an extruded profile were additionally annealed at 300ºC at various times. Hot workability of these various initial states was studied by hot compression tests in the strain rate range 0.01– 10 s-1, the temperature range 300–580ºC up to a strain of 0.9. Additionally tensile tests were carried out to determine the mechanical properties on specimens previously annealed at 300ºC. The oxide surface prevents grains coarsening during hot deformation and consequently excellent mechanical properties of the specimens annealed at a temperature of 300ºC were obtained. Keywords: P/M 1080 Al composite, hot extrusion, annealing, hot compression, mechanical properties.

1 Introduction Demands for weight reduction of parts applied in cars and aircrafts has lead to the replacement of conventional materials with lighter and stronger materials. Aluminum alloys have a great potential for this application but their weaknesses are too low strength and stiffness. On the other hand production of parts from metal-matrix composites which is based on the reduction of grain size is very promising due to their unique mechanical properties in comparison with coarsegrained materials [1–6]. Powder (P/M) and ingot metallurgy which include particle reinforcement mixing with solid or molten metal assure production of composites with WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line) doi:10.2495/MC110211

240 Materials Characterisation V improved mechanical, physical and thermal properties. Hard particles like ceramic, carbides, oxides or nitrides are usually finely dispersed in Al alloy matrix. For P/M the size of the powder can be in the range between several tens of nm to hundred μm, while the sizes of reinforcing particles are accordingly lower. Thus with P/M a finer microstructure with a uniform distribution of micro-constituents can be obtained which leads to mechanical properties, i.e. stiffness, strength and wear resistance, etc. which are usually better than those obtained for materials produced by casting procedures [4–16, 23–25]. Mechanical parts made from aluminum-matrix composites (AMCs) are usually produced by conventional hot deformation processes such as rolling, extrusion, forging, etc. Thus understanding of the behaviour of the material during hot working is essential for optimization of the intrinsic workability, control of evolution of microstructure, and for optimization of the production process. Hot workability studies revealed that AMCs are usually more sensitive to strain rate and temperature in comparison to conventional aluminum alloys. Namely, grain boundary sliding that takes place at higher temperatures and/or lower strain rates can lead to the formation of wedge-shaped cracks at grain boundaries. Additionally, the presence of hard particles in the softer matrix can result in plastic flow localization at the particle-matrix interface. Furthermore during hot extrusion of AMCs prior grain boundaries are usually removed due to diffusion, recovery, etc. that results in reduced mechanical properties. Increasing of grains during processing can be partly prevented by fine dispersion of ceramic particles. For such ultra-fine grained materials with a mean grain size of 1μm and below, different deformation mechanisms are operative than for coarser-grained materials. Namely in very small grains the accumulation of dislocations become difficult that leads to decreased ductility of fine grained AMCs, also at room temperature [5–23]. While for conventionally (by casting) produced AMCs the hot deformability has been studied in several publications [14–22, 26–30], hot deformation of P/M AMCs has found considerably less attention [5, 23–25, 31– 34]. Only a few studies on P/M AMC using pure aluminum as a matrix exist. These studies report that pure Al may benefit more from brittle reinforcement particulates than Al-alloys [5, 35, 36]. The aim of this work is to produce a new fine grained P/M Al-special matrix composite on the base of almost pure Al with an oxide surface on grains aiming to prevent excessive grain coarsening during hot extrusion and followed by a hot deformation process as well as assessment of appropriate hot working parameters. Furthermore the mechanical properties were determined after the extrusion process as well as after various annealing times.

2 Experimental procedures 2.1 Material and characterization of microstructure A JSM-5610 energy dispersive spectrometer from iXRF Systems Inc. with a digital processor 500 for chemical analysis of P/M Al-special was used along



241

with a transmission electron microscopic (TEM, JEOL-2021) for microstructural characterization. 2.2 Processing of initial P/M Al-special The P/M Al-special was produced according to the routes given in fig. 1. Powder was produced by Vacuum Powder Metallurgy, with a special grain powder process. Adding of various additives and lubricants is aimed to accelerate the diffusion processes in bulk material that represents the next step of preparation of powder mixture. After Rubber Isostatic Pressing (RIP) or Cold Isostatic Pressing (CIP) with application of a typical pressure of 100 MPa, the achieved density was 85%. Then follows the thermal preparation of billet (dehumidifying cycle) and machining of the billet to dimensions of 274 mm. The hot extrusion was carried out on an industrial 35 MN indirect press. After extrusion the achieved density was 95% while following hot compression this value was 100% (see fig. 1).

Figure 1:

Processing routes for P/M 1080Al composite.


242 Materials Characterisation V 2.3 Hot extrusion Indirect hot extrusion at various billets temperature and extrusion speeds was employed to convert the processed powder to bulk material. 2.3.1 Extrusion of P/M Al-special on 35 MN indirect press Initial powder billets of dimensions 274 x 1500 mm were hot extruded to a flat bar. Die dimension was 91.3 x 50.8 mm and extrusion ratio was 13.3. Maximum press pressure in the main cylinder is 250 x 105 Pa, which corresponds to a force of 35 MN. Billets 2 and 3 are heated in a die furnace (indirect) to 340ºC and 320ºC, respectively, while billet 1 was heated in an induction furnace to 420ºC. Ram speeds were in the range 3.89–8.87 mm/s. Other data applied at extrusion of P/M Al-special on the 35 MN indirect press are collected in Table 1. Table 1:

Technological data - Flat bar 90x50 (±0.9) mm, alloy P/M Al-special, indirect extrusion.

Technological data; die opening: billet dimension: 274 mm x 1500 mm, 91.3 mm x 50.8 mm, container dimension 280 mm, extrusion ratio: 13.3 Billet material

P/M Al-special 1st billet

2nd billet

3rd billet

Billet length (mm)

500

500

500

Container temperature (ºC)

360

360

360

Billet temperature (ºC)

418

340

320

Ram speed (mm/s)

8.87

6.2

3.89

Puller speed - bar speed (m/min)

7.06

4.94

3.10

Press rest (mm)

50

50

50

2.4 Annealing In order to investigate the influence of elevated temperature on recrystallization resistance, annealing tests were carried out at temperatures of 300ºC and employed at different times, i.e. 1, 5, 9 and 20 hours. 2.5 Hot compression and tensile tests Hot workability was studied by hot compression tests. Cylindrical specimens of Rastegew type with dimensions =10 mm x 15 mm were cut from the extruded profile (3rd billet) so in non-annealed states as well as in annealed states, i.e. previously annealed at 300ºC for 1, 5, 9 and 20 hours, respectively. Hot compression of extruded material was carried out on a computer controlled servo-hydraulic machine, Gleeble 1500D, in the temperature range 300ºC to 570ºC, at five different strain rates (0.01, 0.1, 1 and 5 or 10 s-1) up to a strain of 0.9, while for the annealed material from a production point of view only at a WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


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strain rate range 0.1–5 s-1. The heating rate was 1ºC/s and the soaking time at compression temperature was 1 minute. For reduction of friction between the cylindrical specimen and the tool, and in order to avoid their mutual welding, graphite lubricant and tantalum follies were used. After deformation the specimens were water quenched. The conditions for hot compression testing are given in Table 2. For determination of tensile strength and elongation at room temperature tensile tests were carried out on a Zwick Z400 according to standards (SIST EN 10002-1:2002, 5th edition, June 2002). Tensile specimens was cut out from the centre of the extruded bar, the working length of the tensile specimens was 50 mm with a diameter of =10 mm. Tensile tests were also carried out also on a AA 1050. Table 2: State of material Extruded Annealed

Hot compression test conditions. Hot compression tests Temperature, ºC Strain rate, s-1 300–570 0.01–10 300–570 0.01–10

Strain range 0-0.9 0-0.9

3 Results and discussion 3.1 Composition and microstructures of initial materials The bulk materials consist of 99.44 wt% Al, 0.034 wt% Si, 0.159 wt% Fe, while oxygen (0.363 wt%) is contained on the grain surface. Initial microstructures of loose and compacted powder P/M Al-special are given in figs. 2a-b, respectively. It is visible that the average diameter of powder grains is around 1μm and that the surface layer of grains consists of aluminum oxide (Al2O3) to prevent possible diffusion, recovery processes, i.e. to prevent grain growth during hot deformation.

a

b

Figure 2:

Initial microstructure of P/M Al-special: loose powder (a) and compacted powder with visible oxides on grain surface (b), TEM.


244 Materials Characterisation V 3.2 Indirect hot extrusion of P/M Al-special The time courses of oil pressure in the cylinder for various billet temperatures as well as for ram speeds are shown on fig. 3. The lowest pressure of 215 x 105 Pa (177 bars) was obtained as expected at the highest billet temperature (418ºC) and highest ram speed (8.87 mm/s). On the other hand at the very beginning of the extrusion of the third billet almost the maximal oil pressure was achieved while the highest pressure at the lowest billet temperature (320ºC) and lowest ram speed (3.89 mm/s). The highest oil pressure 215 x 105 Pa (215 bars) was achieved at lowest billet temperature, i.e. at 320ºC.

Figure 3:

Time course of oil pressure in the cylinder during hot extrusion on the indirect press at various ram speed for P/M Al-special, 35 MN indirect press.

3.3 Hot workability of extruded and annealed P/M Al-special Compression tests were carried out for extruded as well as for annealed states. On fig. 4a the typical microstructure of the extruded profile and on fig. 4b the microstructure of deformed samples which have been previously annealed for 20 hours at a temperature of 300ºC are presented. It is visible that the oxide surface on P/M grains is partly damaged but the average diameter of grains is only slightly increased in comparison to the initial state (ca 1μm, see fig. 2b). The figure clearly reveals that the oxide surface prevents grain growth. Obtained flow curves for extruded P/M Al-special are shown on fig. 5. From deformed specimens and the shape of flow curves it can be seen that carrying out the hot deformation at lower strain rates (0.01–0.1 s-1) is not appropriate since on specimens cracks occurred; this results in rapid fall of flow curves in the strain range 0.4–0.5 as can be seen for the strain rate of 0.1 s-1 on fig. 5, left. The behaviour can be ascribed to the usual problem of low-strain rate/hightemperature plasticity of P/M Al-alloys described in several studies such as in [37–41].



Figure 4:

245

Microstructures of extruded P/M Al-special: extruded (a), and forged and annealed (b).

At higher strain rates the flow curves after achieving the maximal value at the very beginning of hot compression, i.e. in the strain range up to 0.03, the values of flow stress begin slightly to fall approaching a steady state (see fig. 5, right). Moreover, at higher strain rates (1–10 s-1) the cracks did not occur that is expressed also by the shape of flow curves.

Figure 5:

Flow curves of P/M Al-special (extruded state) at various temperatures and stain rates.

As mentioned the hot compression tests were carried out also for samples annealed at various times. The shapes of flow curves obtained after various annealing times and for a strain rate of 5 s-1 are similar to those obtained for the extruded state but in general their values decrease up to an annealing time of 5 hours (see fig. 6). It is visible that values of flow curves are considerably lower (about 15%) for annealed states in comparison to non-annealed states. During the extrusion a small percentage of the oxide layer on the grains are torn. This may allow re-arrangement of some grains with orientation changes. During annealing of extruded material the process of part polygonization starts. After one hour at 300°C material became stable since grain growth was blocked by oxides.



Figure 6:

Comparison of flow curves between extruded and at various annealing times for P/M Al-special; strain rate 5 s-1, annealing temperature 300ºC.

3.4 Obtained mechanical properties The general presentation of the obtained mechanical properties, i.e. tensile strength and elongation, of P/M Al-special (indirect extrusion) is shown in fig. 7. The values of tensile strength of around 209 MPa and 205 MPa for the nonannealed and 20 hour annealed state were obtained, respectively. Furthermore, values for elongation for the non-annealed state were around 24% while for the 20 hour annealed samples these values slightly decreased and were around 23%.

Figure 7:

Tension curves σ vs.  for P/M Al-special at various states and times of annealing.

For comparison for AA 1050 the values for strength were around 80 MPa and for elongation were around 38%. Thus around 2.6 times higher values for tensile WIT Transactions on Engineering Sciences, Vol 72, © 2011 WIT Press www.witpress.com, ISSN 1743-3533 (on-line)


247

strength and 1.6 times lower values for elongation for P/M Al-special in comparison to AA 1050 were obtained. From the obtained results for P/M Al-special it can be derived that this material can be applied also at elevated temperatures. It is also worth mentioning that in our case grains consist only of almost pure Al and without particles in the matrix. Thus the achieved mechanical properties are also on a relative high level.

4 Conclusions New P/M Al-special composite made from almost pure Al with an average diameter of 1μm and with oxides on the grain surface was developed. The achieved density after cold isostatic pressing was 84%, after hot extrusion 96% and 100% after hot compression. The P/M Al-special material exhibited a good combination of strength and elongation properties also at samples previously exposed to elevated temperature. The produced P/M Al-special has a great potential to substitute the classical aluminum alloys and steels in the field of an elevated temperature working environment. The obtained values for tensile strength for P/M Al-special alloy are up to 2.6 times higher and for elongation around 1.6 times lower in comparison to values obtained for AA 1050 aluminum alloy produced by classical technology. The tensile strength of samples taken from extruded billet with an extrusion ratio of 13.3 reaches a value of about 209 MPa while achieved elongation was around 24%. The material exhibits good resistance against increasing of grain size at elevated temperature. After 20 hours of annealing the tensile strength slightly decreased and reached the value of 204 MPa. On the other hand the value for elongation remains almost at the same level, i.e. 23%. Extruded material exhibited low hot deformability at strain rates of 0.1 s-1 and lower, while at higher strain rates the compressed samples were crack free.

Acknowledgements The authors gratefully acknowledge Martin Balogand and Karol Iždinský from Slovak Academy Of Sciences, Institute of Materials and Machine, Mechanics Račianska 75, 831 02 Bratislava 3, for their contribution towards analyzing the powder.

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251

Coarsening kinetics of the bimodal  distribution in DS GTD111TM superalloy V. S. K. G. Kelekanjeri1, S. K. Sondhi2, T. Vishwanath1, F. Mastromatteo3 & B. Dasan2 1

GE Global Research Center, Materials Characterization Laboratory, Bangalore, India 2 GE Global Research Center, Materials Research Laboratory, Bangalore, India 3 GE Oil & Gas, Nuovo Pignone, Materials & Processing Engineering, Florence, Italy

Abstract Coarsening of  precipitates in nickel-base superalloys is one of the mechanisms by which creep damage occurs in these alloys. This is brought about by a concomitant increase in the interparticle spacing, which results in faster dislocation movement and therefore, faster creep strain accumulation. Therefore, it is vital to accurately quantify the  coarsening kinetics, which is essential for creep prediction via microstructure-based continuum damage mechanics models. In the present article, we report on coarsening studies of bimodally distributed  precipitates in DS GTD111TM (Trademark of the General Electric Company). The baseline microstructure of GTD111TM consisted of a cuboidal secondary  population and a much finer, spherical tertiary  population with respective mean radii of 266 and 34nm. Long-term aging experiments were conducted on baseline samples at temperatures of T1

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Computational Methods in Multiphase Flow V (WIT Transactions on the Engineering Sciences) (Wit Transactions on Engineering Sciences)

Computational Methods and Experiments in Materials Characterisation III

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Computational Methods in Multiphase Flow VI (Wit Transactions on Engineering Sciences)

Advances in Fluid Mechanics VIII (Wit Transactions on Engineering Sciences)

Physical methods for materials characterisation

Physical methods for materials characterisation

Computational Methods and Experimental Measurements XIV (WIT Transactions on Modelling and Simulation) (Wit Transactions on Modeling and Simulation; Fourteenth ... Methods and Experimental Measurements)

Transactions on Edutainment V

Computational Methods in Earthquake Engineering (Computational Methods in Applied Sciences)

Computational Methods and Experimental Measurements XV (Wit Transactions on Modelling and Simulation)

Monitoring, Simulation, Prevention and Remediation of Dense and Debris Flow III (Wit Transactions on Engineering Sciences)

Transactions on Rough Sets V

Transactions on Computational Systems Biology

Transactions on Computational Systems Biology

Computational Spectroscopy: Methods, Experiments and Applications

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Air Pollution XIX (Transactions on Ecology and the Environment) (Wit Transactions on Ecology and the Environment)

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Composites with Micro- and Nano-Structure: Computational Modeling and Experiments (Computational Methods in Applied Sciences)

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Environmental Toxicology III (Wit Transactions on Ecology and the Enviromnment)

Air Pollution XVII (Wit Transactions on Ecology and the Environment)

Air Pollution XVI (Wit Transactions on Ecology and the Environment)

Computational Materials Chemistry: Methods and Applications

Sustainable Chemistry (Wit Transactions on Ecology and the Environment)

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