Durability of Building Sealants (Rilem Proceedings, 36)

Durability of Building Sealants Sponsors International Union of Testing and Research Laboratories for Materials and ...

Author: Andreas Wolf

314 downloads 1316 Views 5MB Size Report

This content was uploaded by our users and we assume good faith they have the permission to share this book. If you own the copyright to this book and it is wrongfully on our website, we offer a simple DMCA procedure to remove your content from our site. Start by pressing the button below!
Report copyright / DMCA form

DOWNLOAD PDF

Durability of Building Sealants

Sponsors

International Union of Testing and Research Laboratories for Materials and Structures

Building Research Establishment, Ltd.

Tokyo Institute of Technology, Yokohama

Japan Sealant Industry Association

Dow Corning S.A., Belgium

Durability of Building Sealants Proceedings of the International RILEM Symposium on Durability of Building Sealants, Building Research Establishment, Ltd., Garston, UK 6–7 November, 1997

EDITED BY A.T.Wolf Global New Opportunity Development Manager, Sealants Business, Dow Corning S.A., Seneffe, Belgium

London and New York

This edition published 1999 by E & FN Spon 11 New Fetter Lane, London EC4P 4EE Simultaneously published in the USA and Canada by Routledge 29 West 35th Street, New York, NY 10001 E & FN Spon is an imprint of the Taylor & Francis Group This edition published in the Taylor & Francis e-Library, 2005. “To purchase your own copy of this or any of Taylor & Francis or Routledge’s collection of thousands of eBooks please go to www.eBookstore.tandf.co.uk.” © 1999 RILEM All rights reserved. No part of this book may be reprinted or reproduced or utilized in any form or by any electronic, mechanical, or other means, now known or hereafter invented, including photocopying and recording, or in any information storage or retrieval system without permission in writing from the publishers. The publisher makes no representation, express or implied, with regard to the accuracy of the information contained in this book and cannot accept any legal responsibility or liability for any errors or omissions that may be made. Publisher’s note This book has been produced from camera-ready copy supplied by the editors British Library Cataloguing-in-Publication Data A catalogue record for this book is available from the British Library Library of Congress Cataloging-in-Publication Data International RILEM Symposium on Durability of Building Sealants (2nd : 1997 : Garston, Lancashire, England) Durability of building sealants: proceedings of the International RILEM Symposium on Durability of Building Sealants, Building Research Establishment, Ltd., Garston, UK, 6–7 November, 1997/ edited by A.T. Wolf. p. cm.—(RILEM proceedings; 36) Includes bibliographical references and index. ISBN 0-419-24930-3 (Print Edition) 1. Sealing compounds-Congresses. I. Wolf, A.T. (Andreas T.) II. Title. III. Series. TP988.I58 1997 691'.99–dc21 98–49289 CIP ISBN 0-203-62720-2 Master e-book ISBN

ISBN 0-203-63103-X (Adobe eReader Format) ISBN 0-419-24930-3 (Print Edition)

Contents

Members of RILEM Technical Committee 139-DBS, Scientific Committee, Organising Committee, Symposium Secretaries and Peer Reviewers

vii

Preface

xii 1

1

Effects of environmental exposure conditions and sealant composition on silicone sealant properties D.F.BERGSTROMR, D.DASHNERR and H.KRAHNKE

2

Evaluation of the fatigue resistancc of glazing sealants during cure S.R.BLOCK

24

3

Studies into the long-term durability of elastomeric building sealants (Part 2) H.BOLTET and T. BOETTGER

32

4

Further development of a universal test standard for sealant durability and in-service strength A.P.CERRA, W.S.GUTOWSKIS and S. PETINAKIS

50

5

Effects of joint geometry on the engineering properties of elastomeric structural sealants and adhesives W.S.GUTOWSKlL.RUSSELL and P.CHRISTODOULOU

68

6

New opportunities in sealant diagnostics through dynamic solid’s rheometry and micro-specimen testing W.S.GUTOWSKI, L.RUSSELL, A.P.CERRA and S.PETINAKIS

89

7

Building joint movement monitoring and development of laboratory simulation rigs A.R.HUTCHINSONT.G.B. JONES and K.E.ATKINSON

110

8

Effects of early movement on the performance of sealed joints T.G.B. JONESA.R.HUTCHINSON and K.E.ATKINSON

128

9

Case studies on the inspection and re-sealing of failed joint systems M.A. LACASSET, G.B. JONES and D.L.SCOTT

147

vi

10

Effects of degradation factors on the mechanical properties and on the chemical composition of sealants K.OBA

168

11

Recent trends in sealant adhesion tests N.E.SHEPHARD

180

12

Comparative evaluation of silicone sealants exposed to natural and accelerated weathering S.SUGIYAMA

190

13

Test methods assessing the effect of sealants on the aesthetic appearance of buildings A.T.WOLF

201

Index

211

RILEM Technical Committee 139-DBS Members

Active Members A.T.Wolf (chairman) J.Beasley (secretary) T.Böttger H.Bolte C.Folkes V.Gutowski J.Klosowski M.Lacasse K.Oba

Dow Corning, Seneffe, Belgium Building Research Establishment, Garston, United Kingdom University of Leipzig, Leipzig, Germany University of Leipzig, Leipzig, Germany Fosroc Group Development Centre, Birmingham, United Kingdom Commonwealth Scientific and Industrial Research Organisation (CSIRO), Highett, Australia Dow Corning Corporation, Midland, Michigan, USA National Research Council of Canada, Ottawa, Canada Eidgenössische Materialprüfungs- und Forschungsanstalt (EMPA), Duebendorf, Switzerland

Corresponding Members E.Brandt Danish Building Research Institute, Horsholm, Denmark P.-G.Burstroem University of Lund, Sweden T.O’Connor Smith, Hinchman & Grylls Associates, Detroit, Michigan, USA K.Gertis Fraunhofer-Institut für Bauphysik, Stuttgart, Germany S.Hurley Taywood Engineering, Southall, United Kingdom S.Linde Zel-Aaren Innovation, Boras, Sweden J.-Ch.Marechal Centre Scientifique et Technique du Batiment, Grenoble, France

viii

M.Puterman W.Sharman

Israel Institute of Technology, Haifa, Israel Building Research Association of New Zealand (BRANZ), Porirua, New Zealand

RILEM Secretary General M.Brusin

Ecole National Supérieur, Paris, France

Symposium Organisation

Scientific Committee A.T.Wolf (chairman) M.Lacasse K.Oba

Dow Corning, Seneffe, Belgium National Research Council of Canada, Ottawa, Canada Eidgenössische Materialprüfungs- und Forschungsanstalt (EMPA), Duebendorf, Switzerland

Organising Committee A.T.Wolf (chairman) J.Beasley (secretary) M.Lacasse K.Oba M.Gaddes A.Mondair T.Scalisi

Dow Corning, Seneffe, Belgium Building Research Establishment, Garston, United Kingdom National Research Council of Canada, Ottawa, Canada Eidgenössische Materialprüfungs- und Forschungsanstalt (EMPA), Duebendorf, Switzerland Building Research Establishment, Garston, United Kingdom Building Research Establishment, Garston, United Kingdom Dow Corning, Seneffe, Belgium

Symposium Secretaries M.Gaddes A.Mondair T.Scalisi

Building Research Establishment, Garston, United Kingdom Building Research Establishment, Garston, United Kingdom Dow Corning, Seneffe, Belgium

Peer Reviewers

A.Baskaran H.Brockmann A.Buchholz E.Burnett A.Cerra M.Y.L.Chew P.Descamps P.D.Gorman V.Gutowski J.-P.Hautekeer A.R.Hutchinson J.M.Klosowski G.B.Lowe J.L.Margeson J.C.Myers D.H.Nicastro K.Oba T.O’Connor R.M.Paroli

National Research Council of Canada, Ottawa, Ontario, Canada Universität Bielefeld, Fakultät für Chemie, Bielefeld, Germany Geocel Corporation, Elkhart, Indiana, USA The Pennsylvania Housing Research Center, Pennsylvania State University, USA. C.S.I.R.O. Division of Building, Construction & Engineering, Highett, Victoria, Australia School of Building and Estate Management, National University of Singapore, Singapore Dow Corning, Seneffe, Belgium Gorman Moisture Protection, Inc., E1 Paso, Texas, USA C.S.I.R.O. Division of Building, Construction & Engineering, Highett, Victoria, Australia Dow Corning, Seneffe, Belgium Joining Tchnology Research Centre, Oxford Brookes University, Headington, Oxford, United Kingdom Dow Corning Corporation, Midland, Michigan, USA Morton International, Coventry, United Kingdom Ottawa, Ontario, Canada Simpson Gumpertz & Heger Inc., Arlington, Massachusetts, USA Engineering Diagnostics, Houston, Texas, USA Eidgenössische Materialprüfungs- und Forschungsanstalt (EMPA), Duebendorf, Switzerland Smith, Hinchman & Grylls Associates, Detroit, Michigan, USA National Research Council of Canada, Ottawa, Ontario, Canada

xi

L.B.Sandberg N.E.Shephard S.Spindel K.Tanaka R.W.Tock J.Wightman A.T.Wolf

Department for Civil and Environmental Engineering, Michigan Technological University, Houghton, Michigan, USA Dow Corning Corporation, Midland, Michigan, USA D/L Laboratories, New York, New York, USA Laboratory of Engineering Materials, Tokyo Institute of Technology, Nagatsuta, Yokohama, Japan Department of Chemical Engineering, Texas Tech University, Lubbock, Texas, USA Virginia Tech University, Blacksburg, Virginia, USA Dow Corning, Seneffe, Belgium

Preface

This book contains the proceedings of the Second International Symposium on Durability of Building Sealants which was held on November 6–7, 1997, at Garston, England, under the joint auspices of the British Building Research Establishment Ltd. (BRE) and the International Union of Testing and Research Laboratories for Materials and Structures (RILEM) Technical Committee 139DBS (Durability of Building Sealants). The purpose of a building sealant is to seal a joint between various construction components or materials. The sealant thus provides a barrier against the “elements” such as moisture, driving rain, standing or pressurised water, draughts, sand, dust, etc. All building sealants, once installed, are exposed to environmental degradation factors which affect their performance over time and ultimately cause them to fail. Replacing failed sealant joints is time consuming and can be expensive, representing a substantial proportion of the overall building maintenance costs. Specifiers and building owners therefore need to know the predicted service life of a sealant in order to estimate the overall cost associated with the weatherproofing of buildings. The RILEM Technical Committee 139-DBS “Durability of Building Sealants” was inaugurated in July 1991 and has set itself the following objectives: • to review the present knowledge regarding the assessment of the durability of building sealants; • to promote research in this field; and • to make recommendations for suitable experimental research and test methods.

xiii

In particular, one of the technical committee’s goals is to collect experimental data on the natural and accelerated ageing of sealants and to attempt to correlate data obtained according to different ageing regimes using physical performance and chemical analysis methods. The committee hopes to address the following questions with its work: • what chemical and physical changes occur in a sealant as a result of weathering and what effects do these changes have on the cohesive and adhesive properties of the sealant? • how do changes in the cohesive and adhesive properties influence the failure rate and failure type of a sealant in service and in the laboratory? • how well do artificial weathering tests simulate the service conditions a sealant is exposed to? • how can such tests be evaluated and validated as durability tests? While RILEM does not directly sponsor research, its intent is to act as a catalyst for future research by bringing experts within a field together. The Second International Symposium on Durability of Building Sealants brought together architects, engineers and scientists—researchers and practitioners. Their aim was to transfer new ideas, gained from both laboratory research and field studies, to the study of sealant durability and the development of high performance sealants. Lively discussions during the breaks and over British tea stimulated new ideas and initiated further research into the durability of building sealants. Due to the success of the past symposia and as a result of the continued interest in this research topic, the Third International Symposium on Durability of Building and Construction Sealants is currently being organised and will be held on February 2–3, 2000, in Fort Lauderdale, Florida, USA. This volume contains twelve contributions reflecting the wide spectrum of current state-of-the-art research into sealant durability. The symposium papers cover the following topics: • building joint movement monitoring and development of suitable laboratory simulation rigs; • effects of early joint movement on the performance of sealed joints; • test methods for the assessment of the fatigue resistance of sealants and the tenacity of sealant adhesion; • effects of degradation factors on the mechanical properties and the chemical composition of sealants; • correlation of changes in sealant properties produced by laboratory ageing techniques and those observed in naturally weathered samples or under inservice conditions; • prediction of the long-term performance of building sealants from accelerated weathering data.

xiv

The contributions made by the authors help fill the need for scientific and technological reference and provide the underpinnings for the development of an International Standard on Sealant Durability. In closing, I would like to gratefully acknowledge the outstanding quality of the papers contributed by the authors as well as the dedicated efforts of the scientific committee, the organising committee, the peer reviewers, the session chairmen, the staff of the Building Research Establishment (BRE) and of Chapman & Hall who helped to make the symposium and the resulting publication possible. My special thanks go to Maureen Gaddes, Angela Mondair and John Beasley of BRE for their support in the organisation of the symposium and to Tina Scalisi who patiently edited the manuscripts. Andreas T.Wolf Chairman of RILEM TC139-DBS Seneffe, Belgium August 12, 1998

EFFECTS OF ENVIRONMENTAL EXPOSURE CONDITIONS AND SEALANT COMPOSITION ON SILICONE SEALANT PROPERTIES Effects on mechanical properties and apparent cross-link density D.F.BERGSTROM, R.D.DASHNER and R.H.KRAHNKE Dow Corning Corporation, Midland, USA

Abstract Understanding the effects of environmental and chemical composition factors on the long term degradation of silicone sealants allows for better design and use of these materials. The work described herein examines the effects of water, water vapour, UV, heat and ambient ageing conditions on some commercial and lab prepared sealant properties. In addition, designed experiments were conducted to identify the effects of composition factors such as filler, catalyst or adhesion promoter levels on the ageing of oxime silicone sealants. The effects of exposure were evaluated through measurement of changes in strength, modulus, extensibility, hardness and apparent crosslink density. Changes in both sealant composition and environmental factors show effects on ageing rates and behaviours. The results hint at the existence of complex multiple mechanisms of change, sometime occurring with opposing effects. Keywords: Ageing, composition, crosslink, hardening, immersion, sealants, silicones, weathering. 1 Introduction Silicone sealants are well known for their excellent resistance to environmental degradation, especially with exposure to UV and heat. The stability of silicones is a result of the inherent resistance to cleavage of the SiC and SiO bonds. Silicones have the advantage of maintaining high flexibility over a wide temperature range. These factors allow silicones to out-perform most sealant

2 DURABILITY OF BUILDING SEALANTS

materials under a variety of conditions both in long term durability studies and in real service life [1–6]. Silicone sealants, although durable, are not unaffected by long term degradation under harsh conditions. Under certain exposure conditions, gradual changes can be observed in some properties. Studies have shown very gradual changes in modulus and slow embrittlement with ageing of some silicone sealants. Tock [7] examined the effects of various environmental conditions on crosslink density and modulus. Exposure conditions included dry storage and immersion in neutral, acidic and basic aqueous solutions. Aqueous acid immersion (pH500 ** >500 >500 >500 >500 >500 >500 >500 >500

1 3 1 1 1 1 1 1 1

729 615 713 734 773 845 865 708 751

85 90 2 10 2 70 2 2 9

90 86 345 368 304 147 371 428 407

FURTHER DEVELOPMENT OF A UNIVERSAL TEST 57

Sealan Tensile @ tsubstrate RT * combi nation

Tensile @ 88°C

%AF* UTS* %AF UTS (kPa) (kPa)

D1 D2 D3 D4 D5 D6 D7 D8 D9

0 0 0 0 0 0 0 0 0

997 924 903 902 937 924 1023 882 834

0 0 0 0 0 47 0 0 0

671 588 744 524 749 611 716 640 674

Tensile @ -29°C

%AF UTS (kPa) Correlation coefficient 0 1145 0 1279 0 1264 0 1155 0 1165 0 1230 0 1244 0 1268 0 1376 Corelation coefficent

Tensile after immersion

%AF

UTS (kPa)

CSIRO test

%AF

Failure stress (kPa) -0.96

1 20 1 1 1 1 1 40 1

699 543 778 730 773 845 865 708 751

28 78 2 5 8 38 5 20 1

157 79 302 293 254 166 329 255 255 -0.92

* RT=Room temperature; %AF=per cent adhesive failure; UTS=ultimate tensile strength ** Beyond the capacity of the loading system.

Table 3. Results for all structural sealant-substrate combinations (continued) Sealant Tensile @ substrate RT* combi nation

E9 F1 F2 F3 F4 F5 F6

%AF*

UTS* (kPa)

0 100 100 100 100 100 100

653 529 549 535 505 534 434

Tensile @ 88°C

%AF UTS (kPa) 0 100 100 100 100 100 100

502 180 276 535 505 534 434

Tensile@ -29°C

%AF UTS (kPa) 0 100 100 100 100 100 100

312 783 1755 420 311 1146 858


%AF UTS (kPa) 0 47 13 37 8 12 20

702 514 470 521 470 497 431

CSIRO test

%AF

5 100 100 100 100 100 100

Failure stress (kPa) 231 5 10 5 5 5 5


Sealant Tensile @ substrate RT* combi nation

Tensile @ 88°C

Tensile@ -29°C


CSIRO test

%AF*

UTS* (kPa)

%AF UTS (kPa)

%AF UTS (kPa)

%AF UTS (kPa)

%AF

Failure stress (kPa)

F7 F8 F9

100 100 100

524 264 582

524 100 264 100 582 90 Correlation coefficients

2930 263 1828

251 96 316 0.52

100 100 100

100 100 100

3 100 0

5 5 5 -1

* RT=Room temperature; %AF=per cent adhesive failure; UTS=ultimate tensile strength. ** Beyond the capacity of the loading system.

These new results reinforce previous findings in confirming that the CSIRO test is indeed more discriminatory than other standard tests as currently available. It has now been demonstrated that this applies over a wider range of sealant-substrate combinations including weather-sealants. One of the other important findings, as discussed later, is that the CSIRO protocol can be used to determine the value of the structural strength of the sealant at deformation rates corresponding to the daily thermal movement which is important in structural sealant applications. 3.2 Adhesion performance ranking for sealants and substrates The relatively poor discriminatory value of current tests and standards for assessing the adhesion performance of sealant/substrate systems was previously discussed [1–3] by the authors. This is further supported by the new results presented in Tables 3 and 4, which cover a much broader range of sealants, including both structural and weather-sealants based on various chemical systems. The current results indicate that conventional tensile testing under dry conditions at 23°C, 88°C and -29°C is able to identify adhesion problems in a maximum of only two substrates for the structural sealants. For example, in the case of sealant A, poor adhesion was detected only to substrates 6 and 8. Tensile testing at room temperature after seven days of water immersion also at room temperature further revealed some adhesive failure involving this sealant in combination with substrate 1. The CSIRO test however, was able to detect potential adhesion problems with sealant A on a range of other substrates, thereby allowing a ranking of the adhesive behaviour.


Similar observations may be made for the weather-sealants. For example in the case of sealant K, only the CSIRO test was able to discriminate between the adhesion quality to the various substrates. Considering the broad range of sealant types investigated, it is not unreasonable to suggest that the CSIRO test would offer similar advantages when testing most elastomeric sealants. It is also worth noting that the results of testing using the Hockman Cycle (ASTM C714) on the weather-sealants, generally revealed only minor levels of adhesive failure at the edges of the joints, the average being less than 5%. The only combination that showed any significant signs of adhesive failure was 17 and this also performed poorly when assessed using the CSIRO test. Table 4. Results for all weather-sealant-substrate combinations Sealant Tensile @ RT Tensile @ 88°C substrate combin ation

Tensile@ -29°C


CSIRO test

%AF

UTS (kPa)

%AF UTS (kPa)

%AF UTS (kPa)

%AF UTS (kPa)

%AF

Failure stress (kPa)

G1 G7 G10 G8

47 0 72 0

954 1003 1037 1034

27 0 33 0

2732 3097 2962 1547

0 30 3 0

311 474 298 488 0.48

100 100 2 100

5 5 128 5 —1

H1 H7 H10 H8

0 0 0 0

410 370 424 348

0 0 0 100

1493 1480 1439 1088

17 0 10 100

356 340 330 327 -0.50

20 40 100 40

136 108 95 140 -0.80

I1 I7 I10 I8

5 13 52 0

602 624 675 660

0 0 100 83

1267 1273 1391 1358

0 0 0 33

583 449 628 623

80 100 100 60

62 5 5 80 -0.97

J1 J7 J10 J8

0 0 100 0

734 757 773 733

0 0 100 0

240 0 275 30 267 3 280 0 Correlation coefficients 232 17 215 0 229 10 176 100 Correlation coefficients 417 0 326 0 280 0 367 33 Correlation coefficient 365 0 383 0 440 100 424 77

658 535 352 238

1 100 100 100

80 5 5 5

1375 0 1393 0 1407 100 1037 77


Sealant Tensile @ RT Tensile @ substrate 88°C combin ation %AF

UTS (kPa)

%AF UTS (kPa)

K1 K7 K10 K8

0 0 0 0

767 708 827 847

0 0 0 7

L1 L7 L10 L8 M1 M7 M10 M8

0 0 0 0 32 0 0 0

390 382 382 347 617 568 546 567

0 0 5 0 18 5 0 0

Tensile@ -29°C

%AF UTS (kPa)

Correlation coefficient 553 0 592 0 522 0 275 0 Correlation coefficient 174 57 209 0 179 0 174 0 235 32 239 0 198 0 237 0 Correlation coefficient


%AF UTS (kPa)

%AF

CSIRO test

Failure stress (kPa) -1

707 907 1046 1027

0 0 0 0

601 632 554 658

5 10 18 5

348 296 186 352 -1

1484 1311 1533 1472 2216 2304 2169 1649

57 0 0 0 32 0 0 0

257 368 367 349 472 481 442 398

na 32 5 na 10 25 5 15

na 88 71 na 90 86 100 90 -0.87

3.3 Determination of sealant strength using the CSIRO test Conventional tensile testing appears to produce inconsistent correlations between measured joint strength and level of adhesive failure (%AF). Therefore using such results for the engineering design of sealant joints, particularly structural joints, is problematical. Figure 3 plots the %AF versus the failure stress for those standard tensile tests where significant levels of AF have occurred and it appears that the correlation is very poor. The correlation coefficient R2, given in Tables 3 and 4 and Figs 3–5 is a statistical measure of the linear correlation between any two variates, with a value of unity representing a perfect relationship. Values above 0.8 are deemed


Fig. 3. Correlation between %AF and failure stress for tensile testing after water immersion.

(1) highly significant, while values below 0.5 indicate a poor correlation for the particular circumstances of this testing. Figures 4 and 5 contain the same data for the CSIRO test on the structural and weather-sealants respectively. In this case, these results show a significant correlation between these parameters. This enables an estimation of the sealant strength [Ss] by relating the specimen failure stress [Fs] to the level of cohesive failure [%CF], i.e. Table 5 contains examples of this calculation for sealants A, C and D. The value of Ss can also be estimated by extrapolation of the plots in Figures 4 and 5 and determining the y-axis intercept which corresponds to 100% cohesive failure. 3.4 Further interpretation of the sealant strength value derived from the CSIRO test protocol It was observed in previous publications [1,2] that sealant specimens subjected to the CSIRO test failed at stresses below the cohesive strength of the sealant, as determined by standard tensile tests. It is now possible to further interpret this and explain its engineering significance. It is important to remember that the stress imposed on a sealant specimen in the CSIRO test is constant. Due to the visco-elastic nature of the sealants, the


Fig. 4. Correlation between %AF and failure stress for the CSIRO test on structural sealants.

materials deform through the creep process, as schematically illustrated in Figure 6. The average rate of deformation for sealant A in the slow creep zone during the CSIRO test has been estimated and found to be in the range 0.005 to 0. 002 mm/min. Tensile testing of this sealant was also used to determine the relationship between the stress at failure and deformation rate for the range 0.05 to 250 mm/min. The results, as illustrated in Figure 7a, are typical of the stressstrain relationship for polymeric materials, and the datum point for the CSIRO test is also shown to conform to this relationship. Figure 7b contains the same information for sealant B. Based on these results, it is concluded that the estimated sealant strength, as determined by the CSIRO test, is a measure of the structural load-bearing capacity of the sealant under combined mechanical and hydro-thermal stress. Furthermore this strength value has been determined at deformation rates typical of the thermal movements of glass panes (2500 x 2500 mm) bonded to an aluminium frame by structural silicone [4]. It is suggested that a meaningful value for the structural safety factor, n, can be determined for a sealant with 90–100% AF in the CSIRO test, as follows: (2) where S is the standard deviation for the set of tests, and the design stress of 138 kPa (20 psi) is that which is normally assumed for structural glazing applications. For sealant 1 (bead cross-section of 12 x 12 mm) equation (2) and data from Table 3 yield the following value for the safety factor:


Fig. 5. Correlation between %AF and failure stress for the CSIRO test on weathersealants.

(3) Therefore a conservative safety factor in this situation would be about 1.8.

Fig. 6. Creep deformation of Sealant A during the CSIRO test procedure.


Table 5. Estimates of sealant structural strength from CSIRO test protocol

* Determined by extrapolation of the plots in Figures 4 and 5 to determine the y-axis intercept, i.e. 100 % cohesive failure.


Fig. 7. The relationship between the stress at failure (Fmax) and deformation (test) rate for Sealants A and B (sealant bead cross-section 12 x 12mm).

As pointed out in the Introduction, the complete range of specimens is also being subjected to 5000 hours of QUV (artificial UV light and moisture exposure) and longterm natural weathering at a number of sites internationally. The aim of this work is to establish meaningful weathering correlations between all the test methods described in this report and real-time exposure. 4 Conclusions and recommendations 1. . The results presented and discussed in this report provide further evidence that the CSIRO test offers significant advantages over current tests for the performance-ranking of sealant adhesion. 2. . In the current work, the CSIRO test is shown to provide the same levels of discrimination and sensitivity for weather sealants as it does for structural sealants. 3. A method can be developed which allows the results of the CSIRO test to be used to determine the following: • the adhesion performance of sealant-substrate systems simultaneous hydrothermal and mechanical stress; • the structural strength of the sealant at deformation rates closely resembling daily movements in facade joints due to thermal and mechanical movement; and • the structural safety factor under more conservative conditions involving the application of both hydrothermal and mechanical stress.


5 References 1. 2.

3.

4.

5.

6. 7. 8.

9.

10. 11.

12. 13.

14.

Cerra, A.P. (1995) Development of a new adhesion test for silicone sealants, Journal of Testing and Evaluation, Vol. 23, No. 5, pp. 370–376. Cerra, A.P. and Gutowski, W.S. (1996) Performance-based adhesion testing of structural sealants, in: Science and Technology of Building Seals, Sealants, Glazing and Waterproofing: Fifth Volume, ASTM STP 1271, (ed. M.A. Lacasse), American Society for Testing and Materials, Philadelphia, pp. 209–225. Iker, J. and Wolf, A.T. (1990) Comparison of European and U.S. testing and qualification procedures for structural glazing silicone sealants, in: Building Sealants: Materials Properties and Performance, ASTM STP 1069, (ed. T.F. O’Connor), American Society for Testing and Materials, Philadelphia, pp. 67–78. Gutowski, W.S., Lalas, P. and Cerra, A.P. (1996) Structural silicone in curtain walls, in: Science and Technology of Building Seals, Sealants, Glazing and Waterproofing: Fifth Volume, ASTM STP 1271, (ed. M.A. Lacasse), American Society for Testing and Materials, Philadelphia, pp. 97–112. O’Connor, T.F. and Panek, J.R. (1992) Extended laboratory testing for two structural glazing silicone sealants, in: Science and Technology of Building Seals, Sealants, Glazing and Waterproofing: Second Volume, ASTM STP 1200, (ed. J.M. Klosowski), American Society for Testing and Materials, Philadelphia, pp. 1–9. Fedor, G.R. (1992) Usefulness of accelerated test methods for sealant weathering, ibid., pp. 10–28. Gutowski, W.S., Russell, L. and Cerra, A. (1992) New tests for adhesion of silicone sealants, ibid., pp. 87–104. Minkarah, I., Cook, J.P. and Rajagopal, A.S. (1992) Applicability of artificial weathering tests for performance evaluation of elastomeric sealants, ibid., pp. 201–212. Lacasse, M.A. (1994) Advances in test methods to assess the long-term performance of sealants, in: Science and Technology of Building Seals, Sealants, Glazing and Waterproofing: Third Volume, ASTM STP 1254, (ed. J.C. Myers), American Society for Testing and Materials, Philadelphia, pp. 5–20. Descamps, P., Iker, J. and Wolf, A.T. (1994) Methods of predicting the adhesion of silicone sealants to anodised aluminium, ibid., pp. 95–106. Lacasse, M.A. and Paroli, R.M. (1995) Evaluating the durability of sealants, in: Science and Technology of Building Seals, Sealants, Glazing and Waterproofing: Fourth Volume, ASTM STP 1243, (ed. D.H. Nicastro), American Society for Testing and Materials, Philadelphia, pp. 29–48. Beech, J. and Beasley, J. (1995) Effects of natural and artificial weathering on building sealants, ibid., pp. 65–76. Beasley, J.L. and Jenkins, M.R. (1996) The effect of artificial weathering and movement accommodation testing on building sealants, in: Science and Technology of Building Seals, Sealants, Glazing and Waterproofing: Fifth Volume, ASTM STP 1271, (ed. M.A. Lacasse), American Society for Testing and Materials, Philadelphia, pp. 251–265. Beasley, J.L. (1996) The correlation of modulus changes in building sealants after artificial ageing with those that occur after natural weathering, in: Durability of


15. 16. 17.

18.

Building Sealants, RILEM Proceedings 28, (eds. J.C. Beech and A.T. Wolf), E & FN Spon, London, pp. 17–25. Hugener, M. and Hean, S. (1996) Comparison of short-term ageing methods for joint sealants, ibid., pp. 37–47. Hurley, S.A. (1996) The prediction of long-term sealant performance from dynamic accelerated weathering, ibid., pp. 49–62. Bolte, H. and Boettger, T. (1996) Contribution to validation of laboratory test methods for prediction of the durability of building joint sealants, ibid., pp. 91–103. Dixon, WJ. (1965) The up-and-down method for small samples, Journal of the American Statistical Association, Vol. 69, pp. 967–978.

EFFECTS OF JOINT GEOMETRY ON THE ENGINEERING PROPERTIES OF ELASTOMERIC STRUCTURAL SEALANTS AND ADHESIVES The effects of joint geometry W.S.GUTOWSKI and L.RUSSELL CSIRO Building, Construction and Engineering, Melbourne, Victoria, Australia P.CHRISTODOULOU CSIRO Manufacturing Technology, Brisbane, Queensland, Australia

Abstract The effects of joint geometry on the engineering properties of structural silicone sealants are discussed. In particular, the relationship between the joint cross-sectional area at various aspect ratios commonly used in practice (ratio of sealant bite to glueline thickness) of the joint and its mechanical behaviour is examined. Using the principles of the William-Landel-Ferry (WLF) superimposition technique, the effects of deformation rate, temperature and joint cross-sectional area on sealant strength are determined. This is achieved through the generation of master curves for estimating the strength of the structural sealant, based on the traditionally used temperature-deformation rate shift factor (aT) in addition to the geometry-related shift factor (aG) proposed here. Keywords: Aspect ratio, bead dimension, engineering, joint geometry, sealant, structural silicone. 1 Introduction Elastomeric sealants of various chemistry are extensively used by the building and construction industry for structural and weather-sealing applications. One of the major problems faced by structural engineers, designers and sealant manufacturers is how to reliably determine the load-bearing capability of a joint involving an elastomeric adhesive chosen for structural or non-structural bonding of two substrates, e.g. a pane of glass and aluminium frame in a typical building facade application.

EFFECTS OF JOINT GEOMETRY ON THE ENGINEERING PROPERTIES 69

It will be shown in this paper that current procedures, based on world-wide accepted standards, may inadvertently lead to serious under-design of a structure. This, in turn. may lead to unpredictable failures of joints designed on the basis of today’s approach to facade design and the current limited knowledge on the engineering properties of sealants. The authors of this paper provide extensive information on the influence of various service-related factors such as temperature, deformation rate and joint dimensions on the strength of silicone sealants. Based on this, a generalised master curve is developed that allows for the reliable estimation of the anticipated strength of a joint, of any dimension, in a particular real life application, e.g. in a building facade. In this paper, the results are presented for only one, commonly used elastomeric structural sealant. The relationship presented is, however, of universal applicability and this will be substantiated by further results relevant to one- and two-component sealants with different cure chemistries to be published in the near future. 2 Theoretical background of the William-Landel-Ferry (WLF) approach for superimposing temperature and rate effects on material properties The stress-strain curves of any visco-elastic material, e.g. sealant, can be described by a generalised Maxwell model, as expressed by the following Equation [1]: (1) where S= ε= y= M(τ)= M(τ) • d In τ= τ=

stress; constant deformation (test) rate; strain; relaxation distribution function; contribution to the instantaneous tensile modulus in the time time.

span of In τ to In τ+d In τ; and

It is seen from Eqn. (1) that S/ε is a function only of y/ε and thus experimental stress data (e.g. strength at different deformation rates) can be superimposed to give a single curve in the following co-ordinates:


It has been shown by Ferry [2] that the effects of deformation rate and temperature on the strength of a polymer are equivalent and thus can be superimposed by shifting each individual temperature-related curve to a collective one to give a single master curve. The proposed master curve is described [3] by Equation (2): (2)

where T0= T= aT=

arbitrary reference temperature, e.g. glass transition temperature of the actual test temperature at which the stress-strain curve (S versus y) is de temperature shift factor.

polymer; termined; and

The value of the temperature shift factor, aT, can be estimated using WilliamsLandel-Ferry (WLF) equation [3]: (3a)

where T= T0= C1; C2=

test temperature; reference temperature, e.g. glass transition temperature of the polymer; experimental constants.

According to Ferry [3], in the first application of the WLF equation, average values of C1 and C2 were obtained by fitting experimental data to a range of polymers, and were estimated to be 17.44 and 51.6 respectively [4]. It was, however, pointed out by Ferry that the actual variation of C1 and C2 from one polymer to another was too great to permit the use of those constants as “universal” values, as frequently used by many researchers. Ferry pointed out that in a somewhat better approximation, fixed values of C1=8.86 and C2=101.6 can be used in conjunction with a reference temperature T0 which was allowed to be an adjustable parameter but generally fell about 50°C above Tg, giving rise to the Equation (3b) used by the authors of this paper for the estimation of aT: (3b) WLF formula (3b) gives satisfactory results typically in the temperature range Tg < T < Tg+100°C. An example of the master curve construction using the WLF


approach is illustrated in Figure 1. It has been demonstrated for the tensile strength of adhesive joints made using a PET (polyethylene terephthalate) substrate and polybutadiene-styrene rubber, PBSR (Tg=-40°C). In this work we present the proposed extension to the WLF procedure for master curve construction by including the geometrical dimensions of the polymer, subjected to external stress, into the equation. At this stage we do not provide a theoretical explanation of the unusual phenomenon observed on all


Fig. 1. (a) Tensile strength curves of PET/PBSR adhesive joints versus test rate, at various temperature; (b) master curve for tensile strength of PET/PBSR joints versus reduced test rate (ε aT) [5].

types of elastomeric silicone adhesives (as listed in Section 3.1.1) investigated by these authors.


3 Experimental details 3.1 Materials 3.1.1 Structural sealant DC 795, a one-component alkoxy cure structural silicone adhesive, manufactured by Dow Corning, was used and comprehensively analysed in this study. In order to support the validity of the surprising findings discussed in this paper, the authors have also reported (see Figure 5) limited data on a range of other one- and two-component structural silicone adhesives of various cure chemistries, i.e. General Electrics 1200; Dow Corning 995 and 983; Wacker Chemie SG 193 and VP 7460; and Rhone Poulenc Vec 70. 3.1.2 Substrate Material An anodised aluminium substrate provided by Alcan Australia was chosen for this study in order to provide reliable sealant-substrate adhesion. This was a necessary condition required to attain 100% cohesive failure within the sealant under any of the test conditions used in this work. 3.2 Substrate-sealant tensile specimens Before describing the procedure for preparing silicone adhesive/aluminium substrate samples, it is necessary to define some relevant terms. Figure 2 provides an explanation of the terms relating to the geometry of an elastomeric structural adhesive bead in a typical curtain wall application.


Fig. 2. Explanation of the basic terms regarding dimensions of an elastomeric structural adhesive joint in a typical curtain wall application.

Another term used in this investigation is the sealant bead aspect ratio, defined as: aspect ratio=bite/glueline (4) The dimensions of the adhesive beads in the joints investigated in this project and relevant aspect ratios are listed in Table 1. Table 1. Typical sealant dimensions selected for determining cohesive properties Aspect ratio (bite/glueline) 0.67 0.80 0.92 0.95 1.0 1.1 1.2 1.33 1.5 1.6 1.67 1.84 1.87 2.0 2.5 2.67 2.76

Bead size (bite x glueline) (mm2) 8 x 12 8 x 10 12 x 13 20 x 21 6x68x8 10 x 9 12 x 10 8 x 6 16 x 12 12 x 8 16 x 10 20 x 12 24 x 13 28 x 15 12 x 6 16 x 8 20 x 10 20 x 8 16 x 6 36 x 13

Bead cross-section (mm2) 72 80 156 420 36 64 90 120 48 192 96 160 240 312 420 72 128 200 160 96 468


3.0 3.33

18 x 6 20 x 6

108 120

Silicone adhesive specimens with geometries outlined in Table 1 were prepared in accordance with the procedure illustrated in Figure 3. After appropriate filling with the adhesives and tooling, the specimens were left to cure in a conditioned atmosphere (20 ± 2°C; 55 ± 5% RH) for at least six months. In order to dispel any possible allegations regarding biased validity of data presented (e.g. due to incomplete cure of larger size specimens), the authors present a full set of data obtained on specimens cured for six months, as well as those determined on additional specimens stored for a period of approximately eight years prior to testing (see Tables 4 and 5). It is clearly seen from the data presented that material strength at failure exhibits similar dependency on the cross-sectional area of sealant bead, whether the sealant is cured for six months or eight years. 3.3 Test parameters for determining engineering properties of silicone adhesives The parameters varied during the experiments included the material’s temperature and deformation rate. These were selected to approximately cover the range of typical service conditions encountered by silicone adhesives in building facade and are specified in Table 2. Five replicates of each joint specimen were tested in this work.

Fig. 3. Preparation of specimens for determining cohesive properties of silicone adhesives: (a) preparation of an ‘initial specimen box assembly’; (b) filling box assemblies with silicone adhesives.


Table 2. Test parameters for determining cohesive failure criteria of structural silicone adhesives (crosses indicate the selected test points) Temperature (°C)

Test rate (mm/min) 0.05

–20 +20 +40 +80

0.5 X

5

50

X

X

X

X

X

250 X X X

4 Discussion of results All results on strength and other properties of sealants discussed and presented in this work are obtained with entirely 100% cohesive failure within the sealant material during destructive testing. No sealant/substrate delamination was observed in our experiments due to the choice of appropriate substrate. The standard deviation of all results was typically within the range of ±6 to ±8% from the median value reported in all tables and graphs. 4.1 Influence of adhesive bead geometry on joint strength According to current knowledge, in structural applications, the adhesive/sealant bite is designed to provide sufficient load-bearing capacity under negative wind pressure, whilst the glueline (or joint width) is designed to accommodate all movements of the external cladding panels, e.g. those resulting from thermal movements, in relation to the supporting facade frame. The actual joint dimensions will therefore depend on the loads and movements anticipated to occur in service and may vary from 6 x 6 mm for shop front glazing up to 20 x 20 mm or greater for buildings subjected to harsh service conditions such as cyclones and earth tremors. Most current standards recommend that the joint strength and adhesive/sealant movement capability are determined using standard dimensions for the test material, i.e. 10 x 10 mm or 12 x 12 mm. It is universally assumed that the material properties determined from these tests can be used for all structural calculations, whatever the size of the joint. As explained below, this current assumption is erroneous and has to be revised in order to avoid potential problems with the under-design of a curtain wall system. It is shown by the authors of this paper that it is the product of the joint width and bite, in the form of joint cross-section area (see Figure 2), that vitally influences the load-bearing capability of a joint. Work is in progress to verify the


influence of joint surface area on the value of strain at failure which in turn determines the movement capability of the sealant. 4.1.1 Influence of loading mode and bead geometry on adhesive tensile and shear strength, and extension capability In this study, a range of adhesive bead sizes, as listed in Table 3, was tested in shear and tensile mode in order to determine stress-strain characteristics from which the following sealant properties were determined: • • • • •

tensile strength; shear strength; elongation at break for tensile and shear loading mode; Young’s modulus; and shear modulus.

Table 3 also presents the numerical data on the strength and extension at failure for tensile and shear loading. The relationship between adhesive strength and cross-sectional area of the bead is illustrated in Figure 4a. It is seen from the graph that the material’s strength decreases with an increase in the bead crosssection area. It is also noticeable that the shear strength is similar to the tensile strength. The profile of the strength decrease in relation to increasing bead crosssection area is identical for both loading modes, i.e. shear and tensile. Table 3. Tensile and shear strength and extension at failure of DC 795 structural adhesive under tensile and shear loading for different adhesive bead geometries. Joint Aspect ratio dimensions (bite/glueline) (bite x glueline) (mm)

8x6 12 x 6 16 x 6 20 x 6 8x8 12 x 8 16 x 8 20 x 8 8 x 10 12 x 10

1.32 2 2.66 3.33 1 1.5 2 2.5 0.8 1.2

Strain at failure (%) Strength (MPa) Cross-section area (mm2)

Tension

shear

182 143 102 128 185 153 118 138 192 162

318 321 293 305 320 286 283 278 321 303

Tension shear 0.88 0.81 0.72 0.74 0.84 0.87 0.80 0.75 0.83 0.84

0.98 1.02 0.90 0.68 0.98 0.86 0.89 0.74 0.90 0.91

48 72 96 120 64 96 128 160 80 120


Joint Aspect ratio dimensions (bite/glueline) (bite x glueline) (mm)

16 x 10 20 x 10 8 x 12 12 x 12 16 x 12 20 x 12

1.6 2.0 0.67 1 1.33 1.67

Strain at failure (%) Strength (MPa) Cross-section area (mm2)

Tension

shear

Tension

shear

138 107 148 160 151 137

280 265 285 267 258 159

0.79 0.66 0.70 0.77 0.78 0.65

0.79 0.75 0.79 0.77 0.80 0.62

160 200 96 144 192 240

* Test conditions: temperature +20°C, test rate 50 mm/min

The values of Young’s modulus (E) and shear modulus (G) of DC 795 sealant are not reported in this work since these are the subject of another paper currently under preparation. It is, however, noteworthy to state that both of these parameters are simultaneously dependent on: • sealant bead aspect ratio (as can be anticipated); and • cross-sectional area of the sealant bead. Figure 4b illustrates the relationship between extension at failure and bead crosssectional area for the specimens tested in shear and tensile. It is apparent that DC 795 silicone adhesive loaded in shear is capable of extending approximately twice that obtainable under the tensile load, before failure occurs. That approximate ratio (extension at failure in shear to extension at failure in tensile=2:1) seems to apply to the whole range of bead cross-sectional areas investigated in this experiment, i.e. from 36 to 240 mm2. In order to support our claim that the trends observed for DC 795 may apply generally to silicone-based elastomeric adhesives, in Figure 5 we present the results for other structural silicone adhesives investigated in another study (unpublished CSIRO report). It is explicitly shown in Figure 5 that the strength and extension capability of elastomeric silicone adhesives are significantly dependent on the geometry of the joint. The analysis of these results reveals, surprisingly, that the reduction of these properties is related to the increase of the cross-sectional area of the adhesive bead and does not depend on the bead’s aspect ratio as defined by Equation. (4). The above findings have significant practical ramifications. The most important of these is that the adhesive’s strength and elongation at failure, as determined from the standard 6 x 6 mm or 10 x 10 mm tensile specimens must not be directly used for structural/non-structural design purposes, if the


Fig. 4. The relationship between (a) silicone adhesive strength and joint cross-section area; and (b) extension and joint cross-section area, for the specimens tested in tensile and shear. Test conditions: temperature +20°C; test rate 50 mm/min. Data for DC 795 structural silicone adhesive.

dimensions of the ‘actual’ joints are greater than those of the standard test specimen. A good example for illustrating the seriousness of the consequences of such erroneous assumptions for design purposes may be that of sealant 38. It is seen from Figure 5a that the strength of 10 x 10 mm bead is 0.92 MPa, whilst that of a 21 x 21 mm bead is only 0.43 MPa, which is less than 50% of the ‘standard specimen’ strength. Similar caution has to be applied to assuming the movement capability of an elastomeric adhesive. For instance, a 6 x 6 mm bead of DC 795 has 125% extension at failure, whilst a 21 x 21 mm joint fails at only 62% elongation.


Fig. e 5. Relationship between (a) silicone adhesive strength; and (b) extension capability and the cross-section area of an adhesive element (bite x glueline). Test conditions: temperature 20°C; test rate 50 mm/min. Data for various structural silicone adhesives encoded as DC 795; 9; 34; 18; 38; 66.

A general conclusion based on our findings is that, for design purposes, the only properties to be used for estimating the load-bearing capability and movement ability of the joint are those based either on real-size joint testing, or those estimated using master curves as proposed in the next part of this work.


4.2 Development of master curves for estimating the strength of silicone adhesive 4.2.1 General comments The theoretical aspects of master curve development were described earlier. In this section, the procedure discussed will be implemented using experimental data obtained for DC 795 structural adhesive. In section 4.2.3 we will further extend the current WLF theory to accommodate the influence of material geometry on the strength of the joint. The tensile strength data, as required for master curve development, for various joint geometries and tested over a temperature range of -20° to +80°C and at deformation rates of 0.05 to 250 mm/min are listed in Table 4. 4.2.2 Application of WLF theory for strength master curves As explained earlier, the WLF theory enables superposition of temperature and strain (test) rates into a single master curve with the following co-ordinates: Strength versus aTε ε (5) where ε= T= aT=

test (deformation) rate (mm/min); material and test temperature (°C); and is the temperature shift factor (see Eqn. 3).

The value of the shift factor, a,-, as a function of test temperature was calculated with the use of Eqn. (6) using Tg=-35°C for the DC 795 adhesive and using natural logarithm (In) instead of decimal base in further calculations. Figure 6 illustrates the relationship aT versus T. Table 4. Tensile strength (MPa) after six months of DC 795 structural adhesive at various deformation (test) rates, temperatures and joint dimensions. Joint dimension (mm)

6x6 12 x 6

Test rate (mm/min) Temperature (°C) 0.05

0.5

5.0

50

250

+20

+80

-20

+20

+80

+20

+40

0.6

0.5

0.82

0.71

0.64

0.76 (0.72) 0.85

0.8

-20

+20

+80

1.04

0.93

0.78


Joint dimension (mm)

Test rate (mm/min) Temperature (°C) 0.05 +20

18 x 6 10 x 9 12 x 13 24 x 13 36 x 13 14 x 15 28 x 15 20 x 21

0.5 +80

0.46 0.37 0.30 0.30

-20

0.68 0.49

5.0 +20

0.51 0.5

+80

0.51 0.38

+20

50

250

+40

-20

+20

+80

0.85 0.71

0.72 0.64

0.57 0.45

0.65 0.72 0.65(0.66) 0.58 0.54(0.52) 0.49 0.37 0.54 0.47 (0.47) 0.43 (0.43)

* Data in parentheses are for specimens cured for eight years

Fig. 6. Temperature shift factor, aT, for DC 795 silicone adhesive in the temperature range -40° to+100°C.

Using the strength data from Table 4 and relevant values of the temperature shift factor, a,., for each test temperature we are now able to compute reduced data points for the construction of master curves to be presented in the following coordinates: Tensile strength versus In (aT • e) (6) Figure 7 illustrates WLF master curves developed for three sizes of sealant beads, i.e. 6 x 6, 12 x 13 and 24 x 13 mm. It can be seen from Figure 7 that all strength data points processed using WLF theory yield a single master curve relevant to each individual joint size. This enables the simple calculation of the


Fig. 7. Master curves for DC 795 adhesive beads of 6 x 6; 12 x 13 and 24 x 13 mm.

anticipated (and experimentally verifiable) strength of the given size joint within the following range of real service conditions: • temperature: -20 to +80°C; and • deformation rate: 0.05 to 250 mm/min. Since the aim of this work is to further extend the classical WLF theory by including a geometrical factor, we are not presenting analytical formulas for the three master curves in Figure 7, each of which would be applicable to one specific joint size only. 4.2.3 Extended WLF approach involving the geometry of viscoelastic materials It can be seen in Figure 7 that individual master curves for each joint size of DC 795 adhesive are approximately parallel. The same trend is observed for all seven types of structural silicone adhesives investigated by the authors. This confirms that the individual master curves representing strength versus reduced deformation rate In (aT .ε) for various sizes of adhesive joint are offset horizontally in a similar manner for each elastomeric material. Figure 8 illustrates the principle adopted in this work for the development of generalised master curves, which comprises a typical WLF approach in addition to the horizontal shift of each individual size-specific WLF curve, as proposed by the authors of this paper. The generalised GC-WLF master-curve development involves therefore superimposing all individual WLF master curves (describing the strength versus In (aT • ε) for each material’s size) into a single line that will collectively represent the strength of the elastomeric material in relation to the optional


Fig. 8. Explanation of the geometry-related shift factor, aG, in the extended WLF master curve concept.

deformation rate, temperature and the size of the elastomeric element. To achieve this, one of the original WLF curves in Figure 7 has to be chosen as a reference line. All other WLF strength curves will be shifted to this reference line by the above-mentioned geometry-related shift factor aG. To facilitate the task, all individual size-specific curves should ideally be shifted in one direction. The authors therefore chose the results obtained for the smallest joint size (i.e. 6 x 6 mm; joint area A=36 mm2) as the reference set of data points that give rise to the reference line (see Figure 8). For the materials exhibiting a rectilinear relationship between the strength, the service conditions (T, ε) and material’s size, such as DC 795 silicone adhesive, as shown in Figs 7 and 8, the generalised master curve is represented by a line expressed by Eqn.(7): S=mo+m1 In (A)+m2 In (aTe) (7) where S= material or joint strength; intercept; m0= m1;m2= experimentally determined constants dependent on material; ε= strain rate (mm/min); and A= cross-section area (mm2) of the adhesive joint in the plane paral

lel to the tensile force imposed on the joint (see Figure 2).


Any other shaped curve, as may be pertinent to some materials, will be expressed by an appropriate polynomial. The general principles of curve superimposing will remain similar to the protocol described in this paper. Considering the reference line chosen in this work (6 x 6 mm joint; A=36 mm2 area), it is seen from Eqn. (7) as well as from Figure 8 that the generalised x-co-ordinate of the new master curve is expressed in the following logarithmic form: (8) Equation (8) can be expressed as follows: (9) From Eqn. (9), the generalised x-co-ordinate of the new master curve can be expressed in the following general form: x=In (aG • aTe) (10) It can be seen that Eqn. (10) represents a simple extension of the classical WLF approach by an additional shift term, aG, related to the geometry of the tested polymeric material: (11)

where aG= A= m1, m2=

geometry-related shift factor; joint cross-section area; and experimentally determined coefficients in Eqn. (7).

A computer program has been used to estimate the parameters m0, m1 and m2 of the generalised master curve, as described by Eqn. (7). The results are as follows: m1= m2= m3=

1.23 –0.139 0.029

Considering the above, the generalised GC-WLF master curve for estimating the tensile strength, ST, of DC 795 silicone adhesive is as follows: (12) Table 5 lists the set of experimentally determined strength data for DC 795 silicone adhesive, together with the results estimated with the use of the generalised master curve STdescribed by Eqn. (12) and illustrated in Figure 9.


Fig. 9. Generalised master curve for the tensile strength of DC795 silicone adhesive [Eqn. (12)].

All data obtained through the model [Eqn. (12)] were statistically processed to yield very high correlation coefficient R, and standard deviation, S, for the model, i.e.: R= S=

0.968 0.046 MPa

It can be seen from the assessment of the model error in Table 5, that the generalised GC-WLF master curve approach proposed by the authors enables very accurate estimation of the anticipated cohesive strength of DC 795 silicone adhesive in relation to any temperature, test/deformation rate and the size of the polymeric element (e.g. adhesive joint), as analysed in this paper. The average error of the model is ±6.2% with only two points (Nos. 22 and 34) exhibiting approximately 20–25% divergence. 5 Conclusions 1.. It has been shown in this pape er that the tensile strength th of a polym leric element such as silicone adhesive/sealant in the form of a joint in a building facade depends not only on the deformation rate and temperature, as predicted by current theories, but also on the size (cross-section) of the loadbearing element, e.g. adhesive bead. 2.. The current William-Landel-Ferry (WLF) approach on superimposing the effects of deformation rate and temperature on the polymer strength can be


Table 5. Comparison of experimental data (DC 795 strength after six-month cure) with the model estimates.

* Figures in parentheses are the strength of specimens cured for a period of eight years

extended to include the influence of the size (cross-section area) of the loadbearing element, e.g. adhesive bead. 3. The geometry-related shift factor (aG) is additive in a manner similar to that applicable to the superimposed influence of temperature and deformation rate (aT), as originally developed by WLF.


4. The approach proposed by the authors allows for the reliable estimation of the strength of DC 795 silicone adhesive in the following range of conditions: The accuracy of the model is approximately ±6% with the correlation coefficient R=0.968. 5.. The adhesive/sealant strength data for structural design purposes must not be directly used for any other joint sizes than those satisfying the following: (a) joint sizes equal to tested ones; and (b) joint sizes smaller than tested during the sealant appraisal. As seen from Figures 4a, 5a and 7, joint sizes larger than those tested in the appraisal procedure will always yield lower strength than the standard specimen. This may inadvertently lead to an unacceptable reduction of the structural safety factor (see Figure 5a illustrating the influence of joint cross-section area on the actual strength of a range of structural silicone adhesives). 6 References 1 2.

3. 4.

5.

Alfrey, T. (1948) Mechanical Behaviour of High Polymers, Interscience, New York, London, pp. 183–189. Ferry, J.D. (1950) Mechanical properties of substances of high molecular weight. VI. Dispersion in concentrated polymer solutions and its dependence on temperature and concentration, Journal American Chemical Society, Vol. 72, pp. 3746–3752. Ferry, J.D. (1970) Viscoelastic Properties of Polymers, 2nd edn., Wiley, New York, pp. 317–345. Williams, M.L., Landel, R.F. and Ferry, J.D. (1955) The temperature dependence of relaxation mechanisms in amorphous polymers and other glass-forming liquids, Journal American Chemical Society, Vol. 77, pp. 3701–33707. Gent, A.N. (1971) Adhesion of viscoelastic materials to rigid materials. II. Tensile strength of adhesive joints, Journal of Polymer Science. Part A–2, Vol.9, pp. 283–294.

NEW OPPORTUNITIES IN SEALANT DIAGNOSTICS THROUGH DYNAMIC MECHANICAL ANALYSIS AND MICRO —SPECIMEN TESTING Opportunities in sealant diagnostics W.GUTOWSKI, L.RUSSELL, A.CERRA and S.PETINAKIS CSIRO Building, Construction and Engineering, Melbourne, Victoria, Australia

Abstract Establishing the “fitness for purpose” of structural and weatherseal sealants during the service life of a building constitutes one of the major challenges to structural engineers, building owners and regulatory authorities. Currently a range of tedious and costly surveys and tests is required throughout the entire life of these structures, involving the deglazing of a required number of structurally bonded units, to obtain structural certification. The cost and inconvenience caused by such protocols are frequently prohibitive and a deterrent to the broader acceptance of structural glazing in various countries, particularly in Europe, Australia and seismically affected areas. In this paper we analyse new opportunities for significant simplification of some diagnostic procedures. These opportunities are created through the availability of dynamic solids analysers. They require significantly smaller sizes of specimens for the assessment of mechanical properties of materials than the more common mechanical testers and enable the determination of specific properties of materials that can change subsequent to environmental exposure. These “micro-specimens” can further be used for assessing mechanical properties such as Young’s Modulus, tensile strength and elongation capacity in real-life sized joints. Keywords: Dynamic mechanical analysis, elongation, glass transition temperature, mechanical properties, modulus of elasticity, rheometry, sealants, strength.


1 Introduction Assessing the suitability of a sealant requires a determination of the following properties of the joint components: (a). adhesion retention at the sealant/substrate interface; and (b). mechanical/rheological properties of the sealant, viz. • • • •

cohesive strength; extension capability; state-of-cure (as determined by sealant hardness); and modulus of elasticity.

In this work we address the diagnostics of sealant behaviour through the assessment of mechanical properties using traditional mechanical tests, and the novel tools available through dynamic solids rheometry. Sealant/substrate adhesion will be addressed in a future publication, as this work is currently in various stages of experimentation. 2 Dynamic mechanical testing 2.1 Theoretical background Most polymeric materials such as sealants are viscoelastic, so both the viscous and elastic properties of these materials must be measured to understand their rheology and application behaviour. Dynamic mechanical testing enables both of these properties to be measured simultaneously. In a dynamic mechanical test, a small oscillating strain is applied to a sample and the resultant stress is measured. If the sample behaves as an ideal elastic solid, then the resulting stress is proportional to the amplitude of the strain (Hooke’s law), and the stress and strain signals are in phase as shown in Figure 1a. If the sample behaves as an ideal fluid, then the stress is proportional to the strain rate (Newton’s law). In this case, the stress signal is out of phase with the strain signal, leading the strain signal by 90° (Figure 1b). For viscoelastic materials, such as building sealants and adhesives, the phase angle shift δ between the stress and strain signals occurs somewhere between the elastic and viscous extremes, as illustrated in Figure 1c.

NEW OPPORTUNITIES IN SEALANT DIAGNOSTICS THROUGH DYNAMIC 91

The stress signal generated by a viscoelastic material can be separated into two components: an elastic stress that is in phase with strain, and a viscous stress that is in phase with the strain rate. The elastic stress is a measure of the degree to which the material behaves as an elastic solid. The viscous stress is a measure of the degree to which the material behaves as an ideal fluid. By separating the stress into these components, both the strain amplitude and strain rate dependence of a material can be simultaneously measured. 2.1.2 Modulus The elastic and viscous stresses can be related to material properties through the ratio of stress to strain, i.e. modulus. Thus, the ratio of the elastic stress to strain (γ: tensile strain, or τ: shear strain) is referred to as the storage modulus (E’: storage modulus relevant to tensile or compressive mode of loading; G’: storage modulus relevant to shear mode of loading). These moduli (E’ and G’) represent the ability of a material to store energy elastically. The ratio of viscous stress to strain is referred to as the viscous (or loss) modulus; E” in tensile/compression and G” in shear, and is a measure of a material’s ability to dissipate energy. The complex modulus (E=[(E’)2+(E”)2]0.5 or G = [(G’)2+(G”)2]0.5), is a measure of the overall resistance of a material to deformation. The approximations E ≈ E’ and G ≈ G’ can be made when damping is low. The ratio of energy lost (dissipated as heat) per cycle to the energy stored (and hence recovered) per cycle is called the loss tangent: tan δ=E”/E’ or tan δ=G”/G’. Tan δ is a measure of the internal friction of a material and is dimensionless. 2.1.3 Linear viscoelasticity Dynamic measurements are usually made in the linear viscoelastic region of the material, where the dynamic moduli are a function of only temperature and frequency. If the moduli vary with strain, as in the case of building sealants, then the material is said to be nonlinear and its behaviour is far more complex. Many materials are linear only up to some critical strain, deviating from linear behaviour at larger strains. Because of this limitation, precise control of the oscillation amplitude is needed to ensure that experiments are carried out in the linear viscoelastic region.


Fig. 1. The schematic stress response of an elastic, a viscous and a viscoelastic material to a sinusoidally applied strain for the following types of solids: (a) ideal elastic solid (Hookean solid); (b) ideal (Newtonian) liquid; and (c) viscoelastic solid.

2.2 Selected aspects of dynamic mechanical analysis relevant to the assessment of building sealant properties 2.2.1 Dynamic mechanical analysis Dynamic thermal analysis characterises the temperature dependence of the


Fig. 2. Schematic representations of the behaviour of an elastomer at various stages of deformation: (a) under no stress; (b) chain alignment under an applied stress; and (c) stress relief produced by chains slipping past one another into new positions in the sample and recoiling [1].

material’s rheological parameters. For solids, the degree of crystallinity and other morphological features can be examined in this way. These types of tests provide the most sensitive means for measuring the glass transition and other secondary transitions. In amorphous polymers the molecular chains are initially in highly coiled shapes. The application of a stress causes rotation about the chain bonds resulting in an elongation of the molecules in the direction of the stress [1]. This produces a distribution of chain conformations which differs significantly from the most probable distribution and, as this is an unstable state, the chains will rapidly recoil when the stress is released in an attempt to regain their original shape distribution. For short periods of stress the entanglement and intertwining of chains with their neighbours acts as a physical restraint to excessive chain movement and the polymer regains its original length when the stress is removed. If however the stress is maintained for a sufficient time, there is a general tendency for chains to unravel and slip past one another into new positions where the segments can relax and regain a stable coiled form. The resultant flow relieves the tension and produces the observed stress decay. When the temperature increases further and further above the glass transition, the polymer expands thereby creating more room for movement of each chain segment [1]. This enhanced segmental movement promotes stress decay because of the greater ease of chain disentanglement (Figure 2). A typical E’ (storage modulus) versus temperature curve for linear amorphous polymers is shown in Figure 3 and illustrates the various regions of viscoelasticity [2]. I. In the glassy state cooperative molecular motion along the chain is frozen causing the material to respond like an elastic solid to an applied stress. II. In the leathery state (or transition region) the modulus (E’ or G’) drops sharply (in the order of three decades). The glass transition temperature Tg is located in this area. The rapid change in modulus reflects the constant increase in molecular motion as the temperature rises from Tg to about Tg +30°C. Just above the Tg the polymer chains are strongly entangled and the


Fig. 3. Storage modulus as a function of temperature for an amorphous noncrosslinked polymer [2].

movement of the chain segments is still rather slow, imparting leathery type properties to the material. III. In the rubbery state, at temperatures higher than about 30°C above Tg, the modulus (E’ or G’) remains fairly constant with temperature due to an increase in the free volume and molecular motion with temperature and the polymer’s ability to regain its original length once the applied stress is removed. IV. In the viscous state (this applies to thermoplastic materials only; chemically cured/cross-linked sealants typically do not exhibit this type of behaviour) there is little evidence of any elastic recovery in the polymer and all the characteristics of a viscous liquid become evident. In this region, there is a steady decrease of E’ as the temperature increases. As mentioned previously, the complete movement of a chain cannot remain unaffected by the surrounding chains due to the considerable chain entanglement which exists in amorphous polymers. Hence any motion will be retarded by other chains. As a polymeric molecule moves, it will drag along several others and the energy dissipation is then a combination of the friction between the chain plus those which are entangled and the neighbouring chains as they slip past each other [2]. The energy losses during molecular movement are reflected by the loss modulus, E”. The ratio of loss modulus, E”, to storage modulus, E’, defines the so-called “dissipation factor”, i.e. tan δ A typical tan δ versus temperature curve for a linear amorphous polymer below its glass transition temperature is shown in Figure 4 [3]. The peak of tan δ signifies the onset of a new mode of molecular motion with increasing temperature [2]. When a polymeric material is at a temperature below its Tg, large segmental chain motion is frozen. However, other transitions may occur caused by the molecular movement of short sections of the main chain or


Fig. 4. Tan δ as a function of temperature for an amorphous polymer.

of side chains. These transitions are illustrated as small peaks associated with secondary transitions β y etc., in order of descending temperature. A β–transition peak is a measure of the relaxation of side groups attached to the main chain and the magnitude of tan δ at the β peak correlates with material toughness and impact resistance [2]. A y-transition peak is a measure of the glass relaxation of short chain segments within the chain backbone, such as (-CH2-)n sequences where n ≥ 3 or 4. The yrelaxation peak is dependent on density. When the density is increased, the y peak is reduced in magnitude. This is because the magnitude of the peak is a function of the free volume existing in an amorphous polymer [2], As the temperature is increased, the chain segments begin to move and a transition from a glass to a rubber-like state begins to take place [2]. The temperature of maximum damping in this region [2] (denoted by a maximum in tan δ) is usually associated with the glass transition temperature (Tg or a transition). This maximum appears because the polymer is passing from the glassy state where there is no internal friction, through the transition region where the internal friction is high as chain segments begin to mobilise, to the rubber-like state where a larger number of chains begin to move with greater freedom and speed due to an increase in the free volume, hence less internal friction. By analysing marked changes in storage modulus (E’), loss modulus (E”) or loss factor (tan δ) versus temperature (T) curves, the glass transition can be detected as a sudden and considerable change in the property analysed, e.g. the region of the drop in the storage modulus (E’) in the vicinity of an accompanying peak of tan δ or E”. The temperature at which this sudden change occurs is the glass transition temperature Tg (see Figure 4).


2.2.2 Interpretation of the dynamic mechanical curves In dynamic mechanical tests the location of the transition regions and Tg depends on the experimental approach used [2]. The transition region is dependent upon several factors namely: (a). rate of temperature change; (b). frequency; (c). chain flexibility; (d). branching and cross-linking; (e). molecular structure (steric effects); and (f). molecular weight. (a) Rate of temperature change If it is assumed that in the transition region the movement restrictions, still present in the sample, allow only a few chain segments to move in some arbitrary time interval, say 30 seconds, then considerably fewer will have moved, if the observation time is less than 30 seconds [2]. It follows that when a polymer undergoes temperature changes which involve increments of shorter soak times, the chain segments are allowed only a short time to respond to the external load or deformation. Hence, Tg is higher for short temperature soak times than for a similar test with longer soak times (see Figure 5). (b) Frequency

Increasing the frequency (i.e. running a faster test) will increase Tg for the same reasons as mentioned above. Tg is found to increase 5–7°C for every tenfold increase in the frequency [2]. The magnitude of this change is material dependent. (c) Chain flexibility

Chain flexibility is a measure of the ability of a chain to rotate about the constituent chain bonds. Hence, a flexible chain has a low Tg, whereas a rigid chain has a high Tg. Tg can be increased by inserting groups, or by increased cross-linking, which stiffen the chain by impeding rotation, so that thermal energy is required to set the chains in motion. (d) Branching and cross-linking

When cross-links are introduced into a polymer, the density of the sample is increased proportionally. As the density increases, the molecular motion in the sample is restricted and Tg rises. The fall in E’ at Tg is much smaller for polymers with higher degrees of cross-linking because the network structure prevents flow from occurring [2].

NEW OPPORTUNITIES IN SEALANT DIAGNOSTICS THROUGH DYNAMIC 9 77

Fig. 5. Schematic presentation of storage modulus, E’, as a function of temperature, T, for a normal and a fast test [2].

The reverse is evident for polymers with a large degree of branching. The decrease in E’ at Tg is larger for these polymers because the less restricted environment makes it easier for the molecules to flow with increasing temperature. (e) Molecular structure (steric effects)

When the polymer chains are unsymmetrical, with repeat units of the type (CH2CHX), an additional restriction to rotation is imposed by steric effects (i.e., effects of crowding in the molecule). These arise when bulky pendant groups hinder the rotation about the backbone and cause Tg to increase. The size of the side group also affects Tg, and there is some evidence of a correlation between Tg and the molar volume (Vx) of the pendant group [2]. Polar groups also tend to encourage a higher Tg than non-polar groups of similar size due to an increase in the lateral forces in the bulk state hindering molecular motion. (f) Molecular weight

The value of Tg is a function of the polymer chain length. At high molecular weights the glass transition temperature is essentially constant, but decreases as the molecular weight of the sample is lowered. Low molecular weight polymers contain short chain ends which cause less resistance to untangling when a stress is applied. In terms of the free volume


concept, each chain end requires more free volume in which to move about than a segment in the chain interior. With increasing thermal energy the chain ends are able to rotate more readily than the rest of the chain, consequently the glass transition temperature is lowered [2]. 3 Experimental procedures 3.1 Materials and specimen preparations 3.1.1 Sealants Two types of sealants were used in this work: (a). Dow Corning 795, a one-component RTV1, alkoxy cure structural silicone sealant; (b). S–X sealant, a one-component RTV1, unknown cure, structural silicone sealant, manufactured in Australia. 3.1.2 Substrate Clear anodised aluminium was used for specimen preparation in order to achieve good sealant-substrate adhesion always resulting in 100% cohesive failure within the sealants during tensile testing. 3.1.3 Specimen preparation (a) Tensile specimens

Two pieces of substrate, 50 mm long, were used to prepare tensile specimens of varying bead dimensions (glueline x bite): 6 x 6; 8 x 8; 10 x 10; 12 x 12; 20 x 20, 8 x 6, 12 x 6, 18 x 6, 16 x 6, 20 x 6, 12 x 8, 16 x 8, 20 x 8, 8 x 10, 12 x 10, 16 x 10, 20 x 10, 8 x 12, 16 x 12, 20 x 12, 24 x 13, 14 x 15, 28 x 15 mm. Figure 6 illustrates the method used for specimen preparation. (b) Micro-specimens

Micro-specimens for dynamic mechanical analysis and for tensile testing were prepared in accordance with the procedure illustrated in Figure 7. The


Fig. 6. Description and assembly of tensile joints.

dimensions of the sealant beads were as follows: 1 x 1, 1.5 x 1.5, 2 x 2, 3 x 3 and 3 x 3.25 mm. For tensile specimens the sealant was applied with a compressed air gun operating at a pressure of 150–200 kPa which allowed for a controlled flow of sealant into the joint. The tip of the cartridge nozzle was held below the level of the sealant to avoid the entrapment of air bubbles. Joints were slightly overfilled so that a smooth tooled finish could be achieved. In the case of the microspecimens, a small quantity of sealant was applied to the joint and then tooled off. All specimens used in this work were allowed to cure at controlled room conditions (T=20°C; RH=55%) for a period of at least six months. 3.2 Equipment An Instron mechanical tester, equipped with 100 N and 5 kN load cells was used for standard mechanical testing of both types of specimens. An RS A II Dynamic Solid’s Rheometer (Rheometrics) with a 10 N force transducer was used for the dynamic mechanical analysis of micro-specimens.


Fig. 7. Description and assembly of micro-specimens.

4 Results and discussion 4.1 Modulus of elasticity (Young’s modulus) of DC 795 sealant 4.1.1 Static modulus The numerical value of the static Young’s modulus, E=σ/ε, was determined from


Fig. 8. Typical stress-strain curves for viscoelastic materials such as building sealants loaded under tensile or compressive stress.

the linear part of a stress-strain curve, as illustrated in Figure 8. 4.1.2 Dynamic modulus The dynamic modulus of elasticity was determined using tensile specimens shown in Figure 7c in a “temperature step” mode in the range–60 to +200°C, whilst maintaining a constant tensile strain of 10% and frequency of 3 Hz. Other important parameters of the test protocol were as follows: (a) temperature steps

–in the range -60 to -50°C –in the range -50 to -35°C –in the range -35 to +200°C

=5°C =1 °C =5°C

(b) soak time=3 minutes (5 minutes around Tg).

4.1.3 Comparison of Young’s modulus from static and dynamic tests Table 1 contains values of the Young’s modulus for DC 795 sealant for various bead dimensions determined from static stress-strain curves and using dynamic mechanical analysis. The values reported were obtained using three replicates for


Fig. 9. E’, E” and tan δ of DC 795 as a function of temperature (temperature step mode; constant strain—10%; frequency of dynamic strain—3 Hz).

each experimental point. Standard deviation was typically within ±5% of the nominal value shown in the table. The following can be observed from the experimental data shown in Table 1: (a). for test specimens with the same geometry in terms of constant aspect ratios (bite/glueline=1.0), the numerical value of Young’s modulus is similar, whether determined from static stress-strain tests or using dynamic mechanical analysis; and (b). micro-specimens with sealant bead cross-sections (bite x glueline) as small as 1 x 1 mm or preferably 3 x 3 mm can be used for determining the sealant modulus of elasticity instead of the more common tensile test joint specimen. The above observations may have significant practical implications for sealant diagnostics as discussed in Section 5. Table 1. Young’s modulus of DC 795 sealant obtained for various sizes of specimens tested on a Dynamic Solids Analyser RSA II as well as on the Instron mechanical tester.


Aspect ratio (bite/glueline) = 1.0 for all specimens. All moduli are determined for strain=10% Dimensions (bite x glueline) (mm)

Dynamic solid’s analyser RSA II Instron

1x1

1.5 x 1.5

2x2

3x3

3x3

6 x 6 8 x 8 12 x 12

Young’s modulus E (MPa)

0.78

0.79

0.83

0.84

0.82

0.82

0.84

0.82

4.2 Sealant tensile strength and elongation at failure The micro-specimens used in this work exhibit a bead cross-section significantly smaller than sealant beads in real-life applications. This leads to the important question: how are the results obtained with the use of micro-specimens applicable to real engineering solutions and concepts, e.g. curtain walls? In order to provide answers to this, the authors give the following summary of other work that explains the influence of sealant bead geometry on the strength and elongation of the joint. It has been shown by these authors in another work [4] that the mechanical properties, e.g. strength and elongation, of Sealant DC 795 are surprisingly dependent on the joint geometry. Figure 10 illustrates the relationship between the following joint parameters: (a) tensile strength versus joint cross-section area (bite x glueline); and (b) elongation at failure versus cross-section area (bite x glueline). Similar results, to be published in another paper, were obtained for a broad range of one—and two-component sealants of various cure chemistry, i.e. General Electrics 1200; Dow Corning 982 and 983; Wacker Chemie SG 193 and VP 7460; Rhone Poulenc VEC 70. In order to determine any effects due to the state of cure (which may be particularly relevant for the larger specimens), experimental results are presented on specimens cured for six months (open circles in Figure 10) as well as specimens stored for a period of approximately eight years prior to testing (solid circles). It is clearly seen from this data that both strength and elongation at failure exhibit a similar relationship to the cross-sectional area of the sealant bead, regardless of the cure/storage time. It was also shown in another work [4] published in this volume that a relationship exists between sealant strength and the following set of factors: (a) sealant bead crosssection (bite x glueline=36–420 mm2); (b) temperature (–20 to +80°C); and (c) deformation rate (0.05 to 250 mm/min). Equation (1) describes the relationship between tensile strength (ST) of the DC 795 sealant and these parameters: (1)


Fig. 10. Influence of joint cross-sectional area, A (bite x glueline in mm2) on tensile strength, ST, and elongation at failure (data for DC 795 sealant) [4]. Note: open circles— sealant cured for 6 months; and solid circles—sealant cured for 8 years. Data for 8-yearold sealant are shown to provide evidence that all specimens used in the experiments (6– month cure) are fully cross-linked. where: aG=(A)4.79 is the geometry-related shift factor; aT=temperature shift factor; and ε – deformation (test) rate (mm/min); and A=sealant bead cross-section area (mm2).

The numerical value of the temperature shift factor aT is determined from the William-Landel-Ferry equation [5]: (2) Equation (2) gives satisfactory results in the temperature range Tg < T < Tg+100° C. Figure 11 illustrates the generalised master curve [4] for the tensile strength of DC 795 sealant using Equation (1). The use of this master curve (see Figure 11 and Equation (1)) allows a simple estimation of a sealant’s strength for any bead cross-section, temperature (–20 to +80°C) and deformation rate (0.05 to 200 mm/ min) within the joint sizes of 9 to 420 mm2. The model described by Equation (1) and illustrated in Figure 11 yields statistically reliable results, as indicated by a correlation coefficient R=0.968 and low standard deviation S=0.046 MPa.


Fig. 11. Generalised master curve [4] for the tensile strength of DC 795 silicone adhesive [Eqn (1)].

Work is currently in progress on the development of a master curve for describing the extension at failure of DC 795 sealant. 5 Opportunities for sealant diagnostics through the use of micro-specimens and dynamic mechanical analysis The current practice of determining sealant serviceability during periodical material appraisal requires a determination of its mechanical properties (cohesive strength, elongation at failure, hardness) and adhesion at the substrate/sealant interface. This procedure requires deglazing a required number of windows. As well as inconvenience to the building tenants, significant costs are incurred to the owner due to the need of reglazing, tedious specimen preparation and testing. The results presented in this work demonstrate that micro-specimens with sealant beads as small as 3 x 3 x 5 mm can be used for determining sealant mechanical properties. Due to the small size of the sample required, no deglazing is needed. Instead, small pieces of material, approximately 5 x 5 x 5 mm can be cut out from the joint perimeter, and the cavities thus created can be conveniently resealed with a new sealant. Deglazing of a whole unit(s) will only be necessary if these tests indicate a problem with the sealant.


5.1 Proposed diagnostics procedure The proposed procedure involves the following steps: (a). Sealant sampling—approximately 5 x 5 x 5 mm pieces to be cut out from the actual building joint. (b). Micro-specimen preparation (see Fig. 7) in accordance with the following procedure: (i). two aluminium substrates 15 x 15 x 3 mm; one face smooth and perpendicular to the sides in order to assure proper specimen preparation; (ii). rectangular sealant specimen (3 x 3 x 5 mm) cut from the joint sealant; and (iii) adhesive bonding of the above sealant specimen to the aluminium . substrate using one of the following alternatives: very thin “smear” of another silicone sealant; or very thin “smear” of a cyanoacrylate adhesive (e.g. Loctite 406). The latter option requires silicone surfaces to be treated. Prior to bonding own experimentation is recommended to achieve good bond quality between the sealant and the substrate. Pretreatment can be carried out by using either “707” Loctite primer or other proprietary primer, typically recommended for priming non-polar pololefms prior to bonding with cyanoacrylate adhesives (applied by brushing-on); or silicone sealant’s solvent or digester, e.g. “Digesil” (applied by brushing on, followed by water rinse (after 30 seconds residence time) followed by alcohol rinse for fast drying followed by 707 Loctite primer). Note: silicone or cyanoacrylate adhesive is to be applied to the aluminium substrates. (iv) gentle “clamping” of substrates to squeeze excessive adhesive from the . bondline, followed by undisturbed overnight cure. Loctite 406–bonded specimens are ready for testing the next day, whilst siliconebonded specimens need to be left for an additional one week. For larger size specimens (e.g. 5 x 5 x l0 or 6 x 6 x l0 mm), if available, a thicker aluminium substrate is recommended. (c). Assessment of sealant mechanical properties: (i) determination of modulus of elasticity as a function of temperature, using the following non-destructive alternatives: • dynamic mechanical analysis over a temperature range–60 to +200°C (see Section 4.1.2 for details of recommended test protocol). This allows the determination of the sealant Young’s modulus versus temperature and glass transition temperature (Tg);


Fig. 12. Temperature dependence of Young’s modulus for a range of structural silicone sealants.

• “static mechanical testing” with a strain limit of 10% using standard mechanical tester and procedure illustrated in Figure 8. Tests to be run at temperatures of–20, +20, +80 and +200°C. (ii). Mechanical testing of specimens under the following range of conditions: • temperature:–20, +20 and +80°C; • test rate: 0.02, 2.5 and 250 mm/min. Note: all specimens to be conditioned for two hours at the test temperature. 5.2 Sealant diagnostics—an example As described in Section 4.1 and demonstrated in Table 1, the numerical value of sealant modulus can be reliably determined by using either “real life” sized specimens tested on a traditional mechanical tester or by using micro-specimens analysed on a dynamic mechanical analyser. The following is an example of the later technique as applied to S–X sealant obtained from a high-rise building in Australia which was approximately 13 years old. Figure 12 illustrates the relationship of storage modulus, E’, of this material to temperature. The solid curve shows the results obtained on the Rheometrics RSA II Dynamic Analyser (specimen dimensions: 3 x 3 x 5 mm; see Figure 7), whilst


the solid points show the value of E determined using the Instron tester (procedure shown in Figure 8). For comparative purposes, E’ versus T plots of two other common commercial sealants, DC 795 and VEC 70, are also shown in this figure. All RSA II data shown in Figure 12 were obtained using three replicates. The Instron data were obtained using 10 replicates cut away from different locations in the window. The increase in modulus for the S-X sealant at elevated temperatures (80 and 200°C) indicates that the material might have excessively cross-linked during service exposure. This may in turn be indicative of a reduction of the original movement capability due to an embrittlement caused by the loss of chain flexibility. Figure 12 indicates that the other two sealants do not show detrimental changes at higher temperature. An extensive research program currently being carried out by CSIRO in cooperation with members of RILEM TC139–DBS Committee will provide further evidence of the phenomena observed, using real-life sized and microspecimens evaluated with the use of traditional tensile tests and the dynamic solid’s rheometry. Larger size specimens (approximately 6 x 6 mm beads) of S-X sealant from the building facade were mechanically tested under variable temperature and test rates. The results illustrating the influence of these parameters on the strength and elongation at failure are shown in Figure 13, using procedures developed by Karpati [6]. It is commonly accepted that the following performance criteria must be met by structural silicone sealants: (a) design stress–0.138 MPa; and (b) design movement capability–15%. The strength of the S–X sealant determined at the test rate relevant to daily thermal movement rates of window materials is 0.42 MPa. In relation to the design stress of 0.138 MPa (as acceptable in structural glazing), this gives a safety factor of 3.04. In the opinion of the authors of this paper, the low value of this safety factor combined with the trend of the “Karpati plot” in Figure 13 shows the unacceptably low strength of the S–X sealant for structural bonding applications. It is apparent to the authors of this paper that the material investigated in this example may also not exhibit sufficient strength and possibly elongation at the slow deformation rates specific for annual thermal movements of facade components (glass window/aluminium frame). The above observations may necessitate reglazing of the units employing S–X sealant.


Fig. 13. (a) Strength versus deformation rate; and (b) elongation versus deformation rate of sealant S–X.

6 References 1.

Cowie, J.M.G. (1991) Polymers: Chemistry and Physics of Modern Materials, 2nd edition, Blackie Academic and Professional, Glasgow, UK.

2.

Challa, G. (1993) Polymer Chemistry—An Introduction, Ellis Horwood Limited, West Sussex, UK. Rheometric Scientific Inc. (1995) Understanding Rheological Testing: Thermoplastics, information brochure. Gutowski, W.S., Christodoulou, P. and Russell, L. (1999) Effects of joint geometry on the engineering properties of elastomeric structural sealants and adhesives, in Durability of Building Sealants, (ed. A.T.Wolf), Chapman & Hall, London, UK. Ferry, J.D. (1970) Viscoelastic Elastic Properties of Polymers, 2nd edition, John Wiley & Sons, New York, USA. Karpati, K. (1974) Extension cycling of sealants, Proceedings XII PATIPEC Congress, Garmisch-Partenkirchen.

3. 4.

5. 6.

BUILDING JOINT MOVEMENT MONITORING AND DEVELOPMENT OF LABORATORY SIMULATION RIGS Collection of parameters for investigating movement during cure A.R.HUTCHINSON, T.G.B.JONES and K.E.ATKINSON Joining Technology Research Centre, Oxford Brookes University, Oxford UK

Abstract Modern construction methods based on curtain walling systems rely on obtaining an effective seal between the joints. The occurrence of wet-applied sealed joint failure can, in part, be attributed to the movements which occur in the external envelope. Such movements can be both large and rapid if large lightweight cladding panels are involved which are also heavily insulated. The rates and amplitudes of actual joint movements were measured on the predominantly south-facing aspects of two modern office blocks in London. One building was concrete-clad and the other was aluminium-clad. Detailed weather records for the monitoring period were collected so that joint movements could be related to panel surface temperatures, air temperatures, solar gain and precipitation. Experimental cyclic movement rigs were developed which could simulate the measured rates and amplitudes of real joint movements, imposed on International Standards Organisation (ISO) tensile adhesion joints. The design of these rigs was refined to enable ease of joint fabrication ‘in-situ’ and greater control over the test parameters. Preliminary experimental results confirmed the adverse effects, particularly on one-part sealants, of joint movement during cure. Keywords: Aluminium, building, cladding, concrete, joint, movement, movement during cure, sealant.

BUILDING JOINT MOVEMENT MONITORING AND DEVELOPMENT 111

1 Sealants in cladding A large volume of gunnable building sealant is used as weather seals in external curtain walling systems and between lightweight cladding panels. There are several causes of joint movement, both long- and short-term, but the most important shortterm cause is variation in panel temperature. This variation causes building joints to move in relation to panel size, panel materials (and their coefficients of thermal expansion and contraction), the presence or absence of insulation behind the panels, and the method of attachment of the panel (s) to the building frame [1,2]. Amongst the research reported on joint movement monitoring is that by Ryder and Baker [3], Karpati [4,5] and Williams [6]. Joint movement imposes cyclic mechanical strain on the seal which can vary in rate and amplitude depending upon the system factors identified above, and as a function of both daily and seasonal weather patterns. The seals in joints are subjected to constant movement and this cyclic strain may cause degradation of the sealant’s appearance, adhesion and mechanical properties [7]. A number of workers have investigated the effects of joint movement on the properties of laboratory specimens which have reached a significant level of cure, typically after four weeks. However, very few investigators have noted that joint movement can occur as soon as fresh sealant is applied to a joint, whilst it is only partially cured, and that this early movement during cure may be a very significant factor in contributing to premature failure. The problem of movement during cure is exacerbated for the case of single component systems, firstly because some systems take a very long time to reach an adequate level of cure and, secondly, because they cure from the surface of the bead [8]. The disturbance of a seal which is only partially cured can lead to a reduction in both the adhesion and mechanical characteristics of the material [9, 10]. In turn, this may lead to premature failure of the sealant system. Anecdotal evidence indicates that particular problems in Europe are associated with the use of one-component sealant systems on dark-coloured aluminium panels with a predominately southerly aspect, and that these problems are made worse for seals applied in Spring and Autumn. 2 Buildings and joints Structures and buildings comprise many component parts, giving rise to joints or “gaps” between different materials and panels, and around openings. Joints are nearly always in a dynamic state but the actual movement which takes place between joints may be attributed to a number of factors, including: • Ground settlement.


• • • •

Moisture movements in materials, giving rise to shrinkage or swelling. Creep, associated with concrete-framed buildings. Heavy semi-permanent loads, such as equipment and machinery. Wind loading, creating pressure differentials between the front and rear of panels. • Vibration from internal machinery or external transport. • Thermal expansion and contraction of panels and components.

With typical curtain walling systems it is thought that movements due to thermal effects are the most significant. The absolute movement of panels and components due to changes in temperature depends upon the coefficient of thermal expansion of the material (s) involved, their size and mass, and the change in component temperature. In general, large and frequent movements are associated with aluminium and plastics, whilst small and slow movements are associated with brick, concrete and stone [1,2]. The two key movement parameters are therefore amplitude and frequency. There is a large variety of cladding and curtain walling systems. Different design philosophies, materials and methods of construction and attachment are used. The movement potential of the joints between panels therefore depends upon the construction of the building itself, the fixing details for individual panels, the materials involved and the compass orientation with respect to solar gain. With so many different factors both responsible for causing movement and for restraining movement, it is very hard to generalise about typical rates and amounts of movement experienced by sealed joints in buildings. The purpose of collecting data on real buildings was twofold; firstly, to obtain information on the rate and amplitude of joint movement in some typical joints and, secondly, to use this information as input for the development of laboratory movement rigs to simulate field conditions. 3 Buildings selected, joints measured and monitoring techniques used 3.1 Concept Information was collected on the predominantly south-facing aspects of two prestigious modern office buildings in London. A concrete-clad and an aluminium-clad building were chosen to provide a comparison of movement with different constructional materials. Measurements were taken across the joints between typical glazed panels on both buildings.


3.2 Time scale The monitoring period was October 1995 to March 1997. 3.3 Monitoring techniques Movement information across both horizontal and vertical joints was recorded using displacement transducers incorporating linear variable differential transducers (LVDTs) measuring to an accuracy of ±0.02 mm. These were mounted in brackets attached to the surface of one panel, bearing on a target attached close to the edge of the adjacent panel. The brackets and target were bonded to the surfaces of the panels using epoxy resin. Electrical signals were led back to data loggers located within the buildings. Data were collected at 15 minutes intervals between the hours of 06.00 and 22.00, and at 1 hour intervals between 22.00 and 06.00. Additional data were recorded at 15 seconds intervals on the aluminium-clad building for a short time. Glass-fibre insulated J-type thermocouples were taped to the surfaces of the cladding panels to provide a record of surface temperatures. 3.4 Weather records A complete interpretation of the joint movement data requires specific information on incident sun and precipitation. Thus, weather records for the relevant time periods were purchased from the Meteorological Office, issued by the London Weather Centre. These records provided dry bulb temperatures, sunshine hours and rainfall. 3.5 Concrete-clad building The reinforced concrete-framed building was constructed in 1988 using a grid with a 6 m span between columns and 4.3 m floor to ceiling height. The precast concrete cladding panels sit within a perimeter frame and have a typical mass of 3300 kg. A schematic elevation and section of a glazed panel to the bottom and left of the joints monitored is shown in Figure 1. The panels are supported from the bottom to make use of the compressive strength of the concrete. They sit on smaller spandrel panels (attached to the building frame) and the load is transferred through two gravity seats located within 1 m of the centreline of the columns. Two location pins, each situated at the top and bottom of the panel, allow for alignment during assembly of the cladding


Fig. 1. Front elevation of concrete-clad panel.

system. Five wind restraints were additionally attached to the particular panel of interest in the investigation. The gravity seats used are designed to allow for movement of the panel whilst the location pins allow for vertical but not horizontal movement. Wind restraints comprising stainless steel rods are fixed rigidly to the building frame, but due to their nature panel movement can still occur. A two-part polyurethane sealant provided a weather-tight seal between the cladding panels, with sealant beads placed at the front and rear of each panel to provide a double seal. The design width of the joints was around 20–25 mm. It is understood that the building joints were sealed in October 1988. Two joints surrounding the panel shown in Figure 1, located towards the top of the south-east face of the building, were monitored using five transducers. These joints represented a horizontal joint 20–22 mm wide between the top of the panel and the coping panel above, and a vertical joint 8 mm wide to the right of the panel. The panel dimensions were 6.2 m wide x 3.75 m high x 0.2 m thick. 3.6 Aluminium-clad building The steel-framed building was constructed in 1992, based upon a 6 m x 9 m grid with 3.825 m floor to floor heights. Three types of panels were used to clad the building; granite on three elevations, brick on one elevation, and powder-coated aluminium at the corners and set-backs, and on most of the top two floors. The lightweight aluminium panels were mounted in steel “strong back” frames, suspended from the top in two places within 1 m of the centrelines of the columns. A schematic elevation and section of a glazed aluminium panel to the bottom and left of the joints monitored is shown in Figure 2.


Fig. 2. Front elevation of aluminium-clad panel.

Stainless steel sections combined with bolts are used for the support system, with adjustment of the bolts enabling panel alignment during cladding construction. The support system design is rigid and does not accommodate movement. Wind restraints are also used to locate the panel to the building frame; these restraints are not designed to support the dead weight of the panel, but instead restrain the panel from moving if differential wind loads are induced. The wind restraints, comprising 12 mm diameter rods, are attached to the panel and building frame rigidly, but they can deflect if panel movement occurs. The design width of the joints was around 25–30 mm. It is understood that the building was first sealed in 1992 using a two-part polyurethane sealant (the same material as used for the concrete-clad building); it was resealed about 12 months later with a one-part polyurethane sealant. Two joints surrounding the aluminium panel shown in Figure 2, located near the top of the south-facing side of the building, were monitored using two transducers. A further joint on the coping panel above the main panel was also monitored. The sealed joint widths of interest were 25 to 30 mm wide at the exterior face, although the gap between adjacent panels was actually less than this because the panel profiles are stepped at their edges. Measurements were taken primarily of the movement in the vertical joints. The glazed aluminium cladding panel dimensions were 6.0 m wide x 3.8 m high x 0.2 m thick. 4 Theoretical joint movements Cladding panel movement is expected to be primarily a function of panel temperature change as caused by solar gain, or by cooling effects such as wind and precipitation. Theoretical calculations using only air temperatures [e.g.11] are bound to be in error. As panel temperatures increase, the material (s) expands


causing a decrease in joint width (putting a sealed joint into compression). Conversely, if the panel temperature decreases then joint width will increase (putting a sealed joint into tension). If panel movement is related primarily to temperature, the rate of change in panel temperature should be proportional to the rate of joint movement (neglecting friction and non-linear aspects of response). A simplistic prediction of joint movement requires a knowledge of panel temperatures, panel dimensions and a suitable value for the coefficient of linear expansion [e.g.12]. The latter item can be very difficult to assign because of the generally multimaterial nature of real cladding panels. However, such a value can be estimated from measurements of panel movements and temperatures. Suitable values can then be substituted into Equation 1: Expansion=α.∆t.l (1) where α= ∆t= l=

estimated coefficient of linear expansion change in temperature original length

5 Results of joint movement monitoring The data indicate that joint width changes correspond closely with changes in panel temperatures, with the rate of change in panel temperature being proportional to the rate of movement of the joint. There is a clear annual trend which indicates that the greater extremes in panel temperature experienced in Spring and early Summer, and to a lesser extent in the Autumn, are reflected in a greater range, and therefore rate of joint width movement. For both types of building, larger movements were recorded for the vertical joints, corresponding largely to the panels being wider than they were tall. Panel movement was very smooth with no apparent “slip-stick” motion. 5.1 Concrete-clad building The annual pattern of data is shown in Figure 3. The average line illustrates clearly the seasonal movements whilst the maximum and minimum lines show the seasonal variations in total joint movement caused by panel temperature changes. Figure 4 shows a typical Spring day. As the panel temperature warmed up to 12°C from–5°C, the horizontal and vertical joint widths decreased by 0.9 mm and 0.95 mm respectively. For the 8 mm wide vertical joint, the maximum total joint movement recorded was 14%. It is also clear that joint movements exhibit a lag in response to temperature change due to the effects of panel mass, and that


Fig. 3. Annual pattern of joint movement for concrete-clad building.

Fig. 4. Springtime temperature and movement for concrete cladding.

the rate of joint movement is small, The maximum movement rate was 0.008 mm per minute. 5.2 Aluminium-clad building The annual pattern of data is shown in Figure 5 for the vertical joint. For this building the extremes of joint movement occurred in the Spring, corresponding to changes in panel temperature of up to 40°C. The total amount of movement was about three times that of the concrete-clad building and the maximum movement rate was 0.025 mm per minute. Figure 6 shows the same typical Spring day as that for the concrete-clad building shown in Figure 4. As the panel temperature increased from–8°C to 30°


Fig. 5. Annual pattern of joint movement for aluminium-clad building

Fig. 6. Springtime temperature and movement for aluminium cladding.

C, the vertical joint closed up by 3.1 mm; this corresponds to a total joint movement of 12.4% for the 25 mm wide joint. Figure 6 also shows that up to ten small joint movements fall within an overall daily movement, which in turn forms part of an annual cycle. Small rapid movements can be caused by changes from sunny to cloudy conditions during the day, and due to the rapid cooling caused by precipitation. These effects are demonstrated clearly in Figure 7 which depicts a relatively dull day during which there were some sunny periods and a little rain. It is clear that incident sun and precipitation exert a large influence over the movement patterns of relatively lightweight cladding panels.


Fig. 7. Daily pattern of movement for aluminium-clad panel joint.

5.3 Summary On a daily basis, the magnitude and rates of joint movement for the aluminiumclad building were about three times higher than those for the concrete-clad building. Thus the seals between the aluminium panels are far more heavily worked than their counterparts between concrete panels. Relatively small cyclic joint movements occur during the day, which are superimposed upon the overall pattern of daily movements; these in turn overlay a seasonal pattern. The actual data for 1996, together with joint width changes expressed as a percentage of a range of potential joint widths, are collected in Table 1. Table 1. Daily joint movement during Spring, expressed as a % of joint width Type of cladding Joint width variation in 1996 Concrete

1.0 mm

Aluminium

3.1 mm

Hypothetical joint width Movement (%) 6 mm 30 mm 6 mm 30 mm

17 (±8½) 3 (±1½) 52 (±26) 10 (±5)


5.4 Laboratory test parameters Experimental inputs were chosen for laboratory evaluation, based upon this monitoring exercise as well as on figures which were judged to be acceptable by Standards organisations. These were 1 and 10 cycles per day at amplitudes of ±71/2% and ±121/2%, with corresponding imposed temperature ranges of 30°C and 60°C respectively. 10 cycles per day may appear excessive with reference to Figure 7, but was chosen to provide a degree of acceleration. It was also felt to be realistic to make experimental joints with and without closed-cell polyethylene backer rod. 6 Development of laboratory cyclic movement rigs 6.1 Experimental considerations Two movement rigs were developed, known as Mark 1 and Mark 2, which were capable of accommodating tensile adhesion joints of a BS3712/ ISO configuration (Figure 8). The sealant bead in these joints measured 50 x 12 x 12 mm. Such joints were chosen so that the key parameters of “25% modulus” and “extension at peak load” could be obtained following relatively short periods of cyclic movement. One over-riding concept was that the sealed joints had to be able to be cycled from time zero. This implied that the joints needed to be made in the movement rig, with adequate access for gunning and tooling. Among the other rig requirements were the abilities to: • • • •

use a variety of substrate materials, make joints with and without backer rod, vary the rate and amplitude of joint movement, impose temperature/humidity cycles.

To satisfy the last point, both rigs were fabricated from stainless steel and brass materials. 6.2 Engineering principles The test joints were arranged in a linear fashion in the rigs in order to allow sealing and tooling. Both rigs employed the principle of offset cams, within a cam box, on a shaft driven by a motor coupled to a gearbox and controller (Figure 9). Linear motion of the cam box was used to subject the joints to tension


Fig. 8. Tensile adhesion joint.

and compression, by an amount which could be varied depending upon the eccentricity of different cams attached to the drive shaft. The shape of the cams provided a sinusoidal pattern of movement. The rate of movement was varied via the motor controller, allowing between 1 and 96 movement cycles per day. 6.3 Mark 1 rig Eight tensile adhesion joints can be accommodated, four on each side of the cam box. The substrates are 75 long x 12 (or 25 with backing) wide x 6 mm thick. The configuration means that four joints are in extension for the first half camshaft revolution whilst the remaining four are in compression. For the second half of the camshaft revolution, the first four joints are in compression whilst the remaining four are in extension. The open design of the rig enables good accessibility for the sealant gun and tools to both sides of the joints. The rig was used under static laboratory conditions of 21°C and 50% rh only. 6.4 Mark 2 rig Twelve tensile adhesion joints can be accommodated, all located on one side of the cam-box; this means that all joints are in tension or in compression at the same time. The substrates are 90 long x 12 (or 25 with backing) wide x 6 mm thick. This rig was specifically designed to operate within an environmental chamber, with the drive shaft running out through the side of the chamber to the gearbox and motor. All joints were fabricated outside the rig due to its inaccessibility within the environmental chamber. A design involving the use of a sub-assembly jig was adopted which allows newly fabricated joints to be placed within the movement rig without disturbing the joints in any way. A variety of types and thicknesses of substrate can be used. These are clamped to substrate holders, the joints are sealed and tooled, and then the sub-assembly jig is transferred to the movement rig; the substrate holders are then slid into precision slots on the movement rig.


Fig. 9. Mark 1 movement rig.

7 Preliminary experimental work The experimental work described below was conducted using the Mark 1 rig. A more detailed and comprehensive programme of work using both the Mark 1 and Mark 2 rigs is described by Jones et al [13], 7.1 Materials Anodized aluminium substrates 75 x 12 x 6 mm were used; this was sulphuric acid anodized T6063 HE9 material. Two market-leading one-part sealants were used; a silicone and a polyurethane material (in conjunction with the manufacturer’s recommended primer). A two-part polysulphide sealant was also used in conjunction with its recommended primer. Polypropylene cubes (12 mm) were used to control the bead dimensions in the control joints, and polyethylene cubes (12 mm) made from backer rod were used in the movement joints. The same closed-cell polyethylene backer rod was employed in some of the joints where appropriate, with sealing taking place only against non-cut surfaces. 7.2 Experimental programme The effects of the number of movement cycles, amplitude and cure time were investigated. Controls (0 cycles/day), 1 cycle/day and 10 cycles/day were used, with three replicates for the controls and two replicates for the cycled joints; 96 cycles/day were also used to represent an unrealistic but extreme condition for the purpose of comparison only. The static controls, together with the joints subjected to cyclic movement, were tested to destruction after intervals of 1, 2, 3, 5, 7, 9, 18 and 36 days. The ambient conditions used were 21°C and 50% r.h. throughout. The test matrix is shown in Table 2.


Table 2. Matrix of testing for each sealant

Joints subjected to 1 cycle/day actually received a cycle lasting for 2.4 hours, the joints being static for the remainder of the time; this corresponds to a maximum movement rate of 0.04 mm/min. The joints subjected to 10 cycles/day also received cycles lasting for 2.4 hours at the same movement rate. The joints subjected to 96 cycles/day received cycles lasting 15 minutes, corresponding to a maximum movement rate of 0.4 mm/min. Once the required number of cycles had been completed, joints were taken out of the rig and tested to destruction at a rate of 5 mm/min. Each test resulted in a load-extension plot from which the secant modulus in tension at 25% and 100% strain, peak load, extension at peak load, and locus of failure were obtained. 7.3 Results Figures 10 and 11 illustrate the broad range of data trends for the key parameters of 25% modulus and extension at peak load for the 1-part sealants. It is clear that cyclic movement during cure reduces the performance of joints made with onecomponent sealant systems quite significantly. Among the general observations are that: • voids were created in the sealant beads; • the sealant beads were deformed permanently; • a loss of adhesion was exhibited by the polyurethane system (on primed aluminium); premature adhesion failure occurred soon after the commencement of movement at 96 cycles/day; • the polyurethane material performed better when the initial movement was compressive; • significant reductions in modulus and extension were exhibited by both products, broadly in proportion to the rate and magnitude of movement; • the reduction in extension was more significant for the case of joints made with the 1–part polyurethane system; • one cycle per day is sufficient to cause a significant reduction in joint performance. Joints made with the two-part polysulphide sealant were subjected to movement at 1 and 10 cycles/day, since 96 cycles/day was considered to be too severe. For


Fig. 10. Effect of early movement during cure on the performance of tensile adhesion joints made from silicone sealant and anodised aluminium substrates.

Fig. 11. Effect of early movement during cure on the performance of tensile adhesion joints made with polyurethane sealant and anodised aluminium substrates.

these cases the polysulphide material performed much better than the one-part products, with only slight reductions in measured joint extension and modulus. A number of joints made with both one-part and two-part sealants, together with backer rod, were also subjected to the same control and movement cycling


regimes (Figure 12). The presence of backing material slowed down the rate of cure development of the one-part materials, as expected. However the most significant aspect was a marked reduction in the performance of both static and “cycled” joints made with the one-part materials, caused by three-sided adhesion. This led to significant decreases in apparent modulus and extension. It was evident that the low surface tension silicone material exhibited good “adhesion” to the backer rod whilst the higher surface tension polyurethane material displayed some “adhesion”; no real affinity for the polyethylene backer rod was evident with the polysulphide sealant. 7.4 Further work The results from this programme were used to refine and formulate the test parameters to be used in subsequent experiments [13]. These included the further use of backer rod, movement frequencies of 1 and 10 cycles/day, amplitudes of ±7.5% and ±12.5%, the inclusion of cement mortar substrates, and imposed temperature cycles appropriate to the amplitudes of movement. All joints were tested at the age of 21 days following 14 days of cycling and 7 days under static conditions. 8 Key observations and conclusions The monitoring of joint movements in actual building facades proved to be very informative. The rates and amplitudes of movements on the concrete-clad building were relatively small, as expected (Figures 3 and 4); nevertheless, the total joint movement in one of the fairly narrow joints monitored was about 14% (i.e. ±7%). The rates and amplitudes of movement on the aluminium-clad building were rather larger (Figures 5 and 6); however, in the fairly wide joints monitored the total joint movement was 12.4%. In absolute terms, the sealants in the aluminium facade accommodated three times the movement of their concrete facade counterparts and were far more heavily worked in terms of rate. It is useful to note that predictions of joint movement characteristics can be made from estimates of panel temperature extremes, and from a 24 hour monitoring exercise to obtain rates of joint opening and closing. The laboratory movement rigs developed allow a comprehensive investigation of the effects of joint movement during cure as soon as the joints have been tooled. It is clear that cyclic movement during cure reduces the performance of aluminium substrate joints made with two particular 1–part sealants quite dramatically, and that the reduction in modulus and extension is even more pronounced in the presence of backer rod. The effect on joints made with a 2– part polysulphide was however minimal. These effects on performance relate to the speed of cure development, tack and adhesion. The results of a more


Fig. 12. Effect of backer rod on sealed joint performance made with anodised aluminium substrates.

comprehensive investigation and analysis of these factors are covered by the authors elsewhere [13]. 9 Acknowledgements The authors express their appreciation to the UK Engineering and Physical Sciences Research Council (EPSRC) for Grant No. GR/K 51549 awarded by the Built Environment Materials for Better Construction Programme. This project was part-funded and guided by an industrial consortium which included: Bovis Program Management, Fosroc International, Morton International and Taywood Engineering. 10 References 1. 2.

Woolman, R. (1994) Resealing of Buildings—A Guide to Good Practice, (ed. A.R. Hutchinson), Butterworth-Heinemann, Oxford. Hutchinson, A.R., Pagliuca, A. and Woolman R. (1995) Sealing and resealing of joints in buildings, Construction and Building Materials, Vol. 9, No. 6, pp. 379–387.


3.

4. 5.

6.

7.

8.

9.

10. 11.

12.

13.

Ryder, J.F. and Baker, T.A. (1970) The extent and rate of joint movement in modern buildings, in Proceedings of the International Conference on Joint Movement and Design, Brighton, United Kingdom, pp. 1–27. Karpati, K.K. and Sereda, P.J. (1976) Joint movement in pre-cast concrete panel cladding, Journal of Testing and Evaluation, Vol. 4, No. 2, pp. 151–156. Karpati, K.K. and Sereda, P.J. (1976) Measuring the behaviour of expansion joints, Batiment International/Building Research and Practice, Official Journal of CIB, Nov./Dec., pp. 346–355. Williams, M.F., Williams, B.L. and Rouleau, N.W. (1995) Monitoring joint movement in a panelized EIFS building, in Development, Use, and Performance of Exterior Insulation and Finish Systems (EIFS), ASTM STP 1187, (eds. M.F. Williams and R.G.Lampo), American Society for Testing and Materials, Philadelphia, pp. 307–318. Lacasse, M.A., Bryce, J.E. and Margeson, J.C. (1995) Evaluation of cyclic fatigue as a means of assessing the performance of construction joint sealants; polyurethane sealants, in Science and Technology of Building Seals, Sealants, Glazing and Waterproofing: Fifth Volume, ASTM STP 1271, (ed. M.A.Lacasse), American Society for Testing and Materials, Philadelphia, pp. 266–281. Allen, K.W., Hutchinson, A.R. and Pagliuca, A. (1994) A study of the curing of sealants used in building construction, International Journal of Adhesion and Adhesives, Vol. 14, No. 2, pp. 117–122. Margeson, J.C. (1992) The effect of movement during cure on sealant strength development, in Science and Technology of Building Seals, Sealants, Glazing and Waterproofing, ASTM STP 1168, (ed. C.J.Parise), American Society for Testing and Materials, Philadelphia, pp. 22–29. Matsumoto, Y. (1992) The effect of building joint movement on outdoor performance of sealants during their cure, ibid., pp. 30–44. O’Connor, T.F. (1990) Design of sealant joints, in Building Sealants—Materials, Properties, and Performance, ASTM STP 1069, (ed. T. O’Connor), American Society for Testing and Materials, Philadelphia, pp. 141–164. Gutowski, W.S., Lalas, P. and Cerra, A.P. (1996) Structural silicones in curtain walls, in Science and Technology of Building Seals, Sealants, Glazing and Waterproofing: Fifth Volume, ASTM STP 1271, (ed. M.A.Lacasse), American Society for Testing and Materials, Philadelphia, pp. 97–112. Jones, T.G.B., Hutchinson, A.R. and Atkinson, K.E. (1999) Effects of early movement on the performance of sealed joints, in Durability of Building Sealants: Second Volume, RILEM Proceedings, (ed. A.T.Wolf), E & FN Spon, London.

EFFECTS OF EARLY MOVEMENT ON THE PERFORMANCE OF SEALED JOINTS A test procedure T.G.B.JONES, A.R.HUTCHINSON and K.E.ATKINSON Joining Technology Research Centre, Oxford Brookes University, Oxford, UK

Abstract The rates and amplitudes of movements from actual building joints were translated to a novel laboratory-based rig. This rig was used to cycle a variety of tensile adhesion joints made with and without backer rod. The key properties of “25% modulus” and “extension at peak load” were used as the mechanical performance indicators, whilst DMTA measurements were undertaken on bulk sealant samples to monitor the development of cure. In general the one-part sealant systems were affected far more by movement during cure than the two-part products. DMTA measurements indicated that the effect of movement was to actually increase the rate of cure of the one-part systems. However, the overall joint performance was reduced because of voiding and deformation of the sealant bead, as well as a loss of adhesion (which was made worse by the presence of a backer rod). The reductions in modulus and extension were broadly in proportion to the rates and magnitudes of imposed movement. It is suggested that the test procedure represents a rapid method of screening candidate sealant systems (sealant/primer/backer rod) which are being evaluated for applications in joints in curtain walling. Keywords: Aluminium, concrete, curing, mechanical properties, movement during cure, movements rig, sealant joint performance.

EFFECTS OF EARLY MOVEMENT ON THE PERFORMANCE OF SEALED JOINTS 129

1 Introduction 1.1 Movement during cure It is known that cyclic building movement is a major factor in the degradation of sealants in sealed joints [1,2]. It has also been suggested that large and rapid cyclic movement in combination with other degrading agents can have a severe and synergistic degrading effect [3]. All modern polymeric sealants derive their characteristic material properties of strength, modulus and ability to withstand cyclic movement when they “cure” and their molecular chains cross-link [4]. Most sealants, and in particular one-part sealants, can take a relatively long time to cure fully and to develop their full performance properties [5]. In curtain walling, joint movement may occur as soon as the sealant is applied [6]. During this period the sealant may be only partially cured and any disturbance can result in reduced joint performance. This phenomenon has been investigated by a few researchers [2,7] but generally in isolation from other practical construction factors such as realistic movement patterns, joint design details and the presence of backer rod. 1.2 Curing of sealant materials In general one-part sealant materials cure as a result of polymerisation initiated by moisture and oxygen [2]. These agents are atmospheric and therefore the onset of cure is at the moment the sealant is expelled from the sealant gun. Initial moisture contact is with the outer section of the sealant bead and a skin is formed rapidly as a result of cross-Iinking of this outer layer. For the inner section of the bead to cure, moisture and oxygen must diffuse through this skin to regions of uncured sealant material. As the cure of this outer region progresses the thickness of the skin increases and the rate of diffusion of catalysing agents to the inner section of the sealant bead is slowed down. One-part sealant materials therefore cure from the outside in, in a relatively slow manner. Two-part sealant materials cure as a result of reactions between the two components of the sealant, although oxygen and moisture may also be required. Cure is therefore initiated as soon as mixing of the two components commences. Since diffusion is not a major factor for curing of these materials, it is expected that the sealant bead cures homogeneously and in a relatively rapid manner in comparison to one-part sealant materials. DMTA can be used to monitor the rate of cure of sealant materials as the change in molecular mobility and hence glass transition temperature (Tg) can be easily measured at various stages throughout the cure. Changes in the Tg can be


followed from the damping property tan δ as viscoelastic materials will exhibit maximum damping around the glass transition temperature. In order to determine whether curing is affected as a result of cyclic movement during cure, joints can be fabricated and cured under extreme cycling conditions, and the polymeric properties investigated. 1.3 Experimental Approach It is suggested that movement during the early stages of cure could significantly affect two major aspects of the properties of sealant joints: the bulk viscoelastic properties of the sealant, and the mechanical performance of the joint. The latter includes several variables such as choice of substrate, the particular sealant system used, curing times and conditions, and the presence of a backer rod. Additionally, the magnitude and rate of movement could also have a significant effect. The purpose of the investigation reported herein was to derive experimental procedures which mimicked practical reality. Data from measurements of joint movement on real buildings [8] were used to provide appropriate rates and amplitudes of movement for experimental joints in purpose built rigs designed by the authors [8]. In simulating reality more closely, experimental joints were fabricated both with and without backer rod. 2 Experimental 2.1 Materials Standard ISO tensile adhesion joints [8] were fabricated using both aluminium and concrete substrates. Aluminium of grade T6063 HE9 was anodised in accordance with BS 1615. The substrate dimensions were 90 mm long x 12 mm wide x 6 mm thick; substrate dimensions for non-standard foam-backed joints were 90 mm long x 25 mm wide x 6 mm thick. Concrete substrate joints were made from cement mortar blocks 90 mm x 25 mm x 12 mm. All aluminium substrates were degreased thoroughly by wiping the surface with a paper towel soaked in 1,1,1 Trichloroethane (“Genklene”) . All concrete substrate surfaces were prepared using carborundum powder after which they were rinsed in water and left to dry thoroughly. Four market leading sealants typically used in the construction industry were chosen for this study: a one-part silicone sealant (SIL), a one-part polyurethane sealant (PU), a two-part polysulphide sealant (2PS) and a two-part silicon-


modified polyether sealant (2SMP). The appropriate primers as recommended by each sealant supplier were used for each sealant/substrate combination. Rectangular section backing material when used was expanded closed cell polyethylene foam. The foam was cut 25% oversize in thickness (i.e. 15 mm thick) to allow for extension within the joint. The pieces of foam used were approximately 75/90 mm long x 13 mm wide x 15 mm thick for the back of the joint and 12 mm long x 13 mm wide x 15 mm thick for the sides of the joint. Sealant was gunned only against non-cut foam surfaces along the length of the joints. Polypropylene blocks were used as spacers for controlling joint length and width on control joints (2 per joint); each block was 12 mm x 12 mm x 12 mm. These were also degreased thoroughly by wiping the surface with a paper towel soaked in “Genklene”. 2.2 Test procedures 2.2.1 Dynamic mechanical thermal analysis (DMTA) evaluation Joints were fabricated without the use of a primer as it was the bulk viscoelastic properties of the sealant that were of interest and not the level of substrate-sealant adhesion. Half the joints were cycled at an amplitude of ±25% for 10 cycles/day at 21°C ±3°C and 50% rh; the others were left under static conditions under a similar environment and used as the control samples. Samples were prepared for these joints and tested at various stages throughout the cure. DMTA analysis was carried out using Polymer Laboratories’ equipment. The system was fitted with a liquid nitrogen delivery system which enabled testing to be carried out at subzero temperatures. Testing was carried out using the shear sandwich arrangement shown in Figure 1, at a frequency of 1 Hz with a nominal peak to peak displacement of 23 µm. A temperature scan was conducted over the range–80°C to +40°C. Values of tan δ were recorded. Samples of uncured sealant were tested together with sections from inner and outer sections of the sealant bead (prepared as shown in Figure 2) after 1, 3, and 7 days of cure. Specimens were prepared immediately before testing to minimise the additional curing of regions of the specimen not previously exposed to the atmosphere. The sample size for testing under shear was a disc 2–3 mm thick of diameter 7 mm.


Fig. 1. Clamping arrangement for DMTA testing in shear.

Fig. 2. DMTA sample preparation.

2.2.2 Mechanical joint testing The data collected from real building movement presented elsewhere [8] showed that typical amplitudes of movement could be ±7.5% and ±12.5% of joint width for aluminium and concrete cladding panels. The ranges of panel temperature fluctuations associated with these amplitudes were approximately 30°C and 60°C respectively. The frequency of cycles experienced by these real panels was highly dependent on the weather conditions as described [8]. Generally one large cycle was experienced over the 24 hour period, and smaller cycles (up to a maximum of 10) were detected on the aluminium cladding panels. Therefore the parameters chosen to input into the movement rig were a combination of temperature cycling over the relevant amplitudes, i.e. at ±7.5% amplitude a temperature range of 30°C and at ±12.5% amplitude a temperature range of 60°C. Within these ranges the joints experienced maximum extension at the lowest temperatures (5°C at +7.5% and–10°C at +12.5%) and maximum compression at the highest temperatures (35°C at–7.5% and 50°C at–12.5%). Both 1 and 10 cycles a day were imposed on these joints and specimens were also cycled mechanically but without the imposed temperature cycling. Control joints which were not subjected to mechanical cycling were fabricated and tested for comparative purposes. Since backing rod is generally used in sealing of joints in


Table 1. Test matrix for joint cycling, showing numbers of replicates used

p20001a0cg110001 *Both with and without backing

buildings, joints were fabricated using the closed cell polyethylene backer rod and cycled under the most extreme conditions, i.e. ±12.5% amplitude at 10 cycles/day with a 60°C temperature range per cycle imposed. The test matrix outlining these conditions is shown in Table 1. The bead width of the control joints was achieved using the polypropylene spacers, and backer rod applied where necessary. The sealant was gunned into the joint and “tooled off” carefully in order to push the sealant fully into the joint against the backing rod and also to produce a flush bead finish, thereby maintaining the required ISO bead dimensions. The static control joints, together with the joints subjected to cyclic movement using the Mark 2 movement rig described elsewhere [8], were tested to destruction after 21 days cure. All joints were subjected to movement during the first 14 days of cure, and then allowed to rest in a static position for the remaining 7 days, prior to testing. Once a joint had completed its cure schedule, tensile testing was carried out immediately using a universal tensile testing machine which subjected the joint to a tensile force at a rate of 5 mm/min. Each tensile test resulted in a load against extension plot. From each graph, secant modulus in tension at 25% and 100% strain, peak load, and extension at peak load were obtained. The locus of failure in each joint was also determined after each test.


Fig. 3. Curing of a typical one-part polyurethane sealant.

3 DMTA analysis 3.1 One-part sealant materials A typical tan δ curve for an uncured and a fully cured one-part sealant is shown in Figure 3. Curing can be followed by observing the shift of the primary peak to the higher temperature (i.e. to the right of the tan δ versus temperature trace), accompanied by a decrease in the height of the very broad secondary transition. Comparisons between the inner and outer sections of the sealant bead at various stages throughout the cure can be seen in Figure 4. The outer section of the bead cures at a more rapid rate than the inner section of the bead, due to the diffusion of atmospheric agents through to the centre of the bead from its edge. The skin which forms on the outside of the bead acts as a barrier to the penetration of moisture and therefore the bead takes several days to fully crosslink. The effects on the tan δ curve of curing under cyclic and static conditions, for both the inner and the outer sections of the sealant bead over the first 7 days of cure, are shown. Imposing movement during the initial stages of the curing process appears to accelerate the rate of cure of the bead. This is observed in the tan δ traces by a decreased height of the secondary transition for the specimens which have been cycled. This effect is apparent for the specimens taken from the inner section of the sealant bead over 1, 3 and 7 days of cure; however, this was only detected in the outer section of the bead after 1 day of cure. After 1 day, the outer section of the sealant bead shows very little change in the height of the secondary transition. This would suggest that the reaction mechanism responsible for the secondary transition is completed in the outer sections of the bead before 3 days of cure. The inner section however has not


Fig. 4. Effect of movement on bead sections of one-part sealant at various stages throughout the cure.

reached this limit at this stage. If cycling was to affect the mechanism of cure, the shape of the tan δ peak would be different for the outer section of the bead after 7 days of cure since at this stage the secondary transition peak has stabilised at its minimum value. It can therefore be said that movement during cure increases the rate of cure for a one-part sealant material but has little effect on the actual mechanism of cure itself. 3.2 Two-part sealant materials Curing of a typical two-part sealant material can also be followed by comparing the tan δ traces of the uncured and the fully cured sealant as shown in Figure 5. The uncured sealant exhibits a very broad transition over the range–40 to +40°C. Curing is accompanied by a decrease in this peak height and the formation of a relatively narrow peak at approximately–30°C. Figure 6 shows a selection of the tan δ curves for both cycled and static sealant beads at various stages throughout the first 7 days of the cure. During the curing process a primary peak is formed which has a side peak as shown in the tan δ trace for 1 day of cure of the outer section of the sealant bead. This secondary peak is then incorporated into the primary peak.


Fig. 5 .Curing of a typical two-part polysulphide sealant.

There appears to be very little difference between the properties of the inner and the outer sections of the sealant bead as determined by the temperature of the maximum tan δ peak. This is to be expected since a two-part sealant cures homogeneously throughout the sealant bead. The shapes of the curves and position of the peaks over 1 and 3 days are unaffected as a result of movement during cure. It can therefore be assumed that the actual mechanism of curing is unaffected by imposing a cyclic regime on the joint. This would imply that movement during cure of the two-part sealants has no effect on the mechanism of cure, or on the cure rate as shown over a 7 day cure cycle. 4 Mechanical joint performance The effects of both mechanical and temperature cycling on joint performance are presented as plots of cyclic amplitude as a function of 25% modulus (E25) and extension at peak load. Data from the static cured joint are presented for comparative purposes. 4.1 Joints made with one-part sealants Figure 7 shows the effect of amplitude, frequency, temperature and backing rod on the performance of silicone/aluminium joints. E25 is decreased as a result of cycling the joint irrespective of frequency, amplitude or temperature. At both amplitudes this decrease was more significant for joints which had undergone temperature cycling. Cycling at 10 cycles/day had more of an effect on the decrease in E25 than cycling at 1 cycle/day. Changing the amplitude from ±7.5% to ±12.5% did not appear to significantly affect E25 for joints which had only been subjected to movement cycling and not temperature cycling as well.


Fig. 6. Effect of movement on bead section of two-part sealant at various stages throughout the cure.

Temperature cycling along with an increase in amplitude resulted in a dramatic decrease in E25. This decrease was comparable to that between the static joints and those which had been both movement and temperature cycled at ±7.5%. The extension at peak load shows similar trends to those observed for E25 i.e. a decrease in joint performance, irrespective of the amplitude, temperature or frequency. One notable difference between the values of extension at peak load and E25 was the change as a result of increased amplitude. Joints cycled at constant temperature, and joints subjected to temperature as well as movement cycling, exhibited very little difference between ±7.5% and ±12.5% amplitudes. This suggests that only a relatively small amplitude is detrimental to joint extension capabilities. Failure in these joints was 100% cohesive for all specimens tested. Figure 8 shows the effect of the same parameters on silicone/concrete joints. It is important to note that the static modulus and extension values measured for the silicone sealant were different for the aluminium and the concrete substrates. The decrease in both E25 and extension at peak load for concrete joints as a result of any form of cycling is apparent. However, the properties as a result of 1 or 10 cycles at constant temperature were almost identical, suggesting that maximum joint deterioration occurs with minimal amplitude and number of cycles. For joints also subjected to temperature cycling there was a difference between 1 and 10 cycles/day; however, there was no change in extension at peak load as a result of increasing the amplitude. E25 decreased as a result of increased amplitude, i.e. the joint stiffness decreased as the amplitude of cycling increased and


Fig. 7. Effect of amplitude, frequency, temperature and backing rod on Silicone/ Aluminium joints. All joints remained at 50% rh for 3 weeks. “Temp” denotes temperature cycling range for given amplitude. “Backing” joints subjected to 10 cycles/ day with temperature changes.

temperature cycling was imposed. Failure in these concrete substrate joints was 50% cohesive and 50% thin film failure ( 2–PU > 2–PS > 1–PU > 1–Si (1) When the sealant is simultaneously exposed to heat, the effect of water, however, may be different. The influence of hot water on the elastic recovery of polysulfide sealants was minor compared to that on polyurethane and silicone sealants [23]. The type of curing agent may play a role in water induced ageing. Dichromate cured polysulfide sealants had superior resistance to swelling in hot water compared with those cured with manganese dioxide [14]. For manganese dioxide cured sealants, acidic degradation products were formed by the immersion in hot water. The water sensitivity of polyurethane sealants is well known. The typical mode of failure, however, is not cohesive in nature. Rather, the increase in strength and modulus, caused by the hardening of the sealant, places additional stress on the bond line and results in adhesive failure [5], According to Myers [25], an extended period of contact to moisture can lead to surface softening of polyurethane sealants before completion of cure. The softening of polyurethane sealants can be also influenced by the type of substrate. According to Beech [22], the softening which occurred upon water immersion was less with cement mortar than with aluminium substrates. Changes in the chemical composition of polyurethane sealants caused by water immersion were determined using FTIR spectroscopy (showing an increase in hydroxyl absorption) [12]. This finding can

EFFECTS OF DEGRADATION FACTORS ON THE MECHANICAL PROPERTIES 175

be interpreted as a hydrolysis of the urethane crosslinks. The phenomena was most pronounced for the product which contained no calcium carbonate filler. Water immersion also was found to affect the FTIR spectrum of a silicone sealant [10]. For silicone sealants, water immersion causes a decrease in tensile strength [26]. However, according to Tock [27], moisture appeared to have little effect on the mechanical properties of silicone sealants, except when strong sulphuric acid solutions (pH Oxime >> Alkoxy (2)


4.3 UV-light Photodegradation of a polymer is usually initiated by UV-light [16]. The degradation by UV-light generally results in hardening of the sealant, caused by a decrease in the average molecular weight between crosslinks. The energy of the UV light induces chain scission, followed by the formation of new crosslinks. Polysulfide sealants generally respond to accelerated ageing tests which include exposure to UV light with surface crazing, increased rigidity, increased tensile strength and increased hardness [30]. When weathered on the straincycling rack, a two-part polysulfide sealant was found to form cavities, while the same material, weathered without cycling, did not show any permanent deformation [2]. Compared to polyurethane sealants, polysulfide sealants are more sensitive to QUV ageing with regard to its effect on their elastic recovery, however, they may perform better under heat (wet or dry) [23]. For polyurethane sealants, the rate of surface crazing caused by weathering was related to the amount of mechanical strain [25]. Cohesive failure close to the substrate was also identified as a failure mode in polyurethane sealants exposed to QUV accelerated ageing, even without the presence of any joint movement [31]. For silicone sealants, an increase in tensile strength was observed after exposure to UV light (in air) [26]. 4.4 Ozone There are only a few papers dealing with the degradation effect of ozone on the mechanical properties of building sealants. Keshavaraj et al. [32] have studied the oxidative effects of ozone on the ageing of silicone sealants. The study showed that the modulus decreased during an initial period due to the initial degradation reaction, however, over extended exposure time, an increase in modulus caused by accelerated cross-linking reactions was observed. During the initial period, some surface cracking was observed. Ozone and moisture acted synergistically only when the moisture was acidic. Under these conditions, the high modulus sealant developed cracks at the border of the regions of differing crosslink density. The medium modulus silicone sealants, on the other hand, showed very good resistance to ozone attack [32].


5 Recommendation for further studies This paper reviewed the effects of degradation factors on the mechanical properties and the chemical properties of sealants. Numerous papers have been published dealing with the effects of ageing under static conditions (specimens are not exposed to any movement). However, there are only a few papers presenting information on the ageing of sealants under dynamic conditions (specimens exposed to cyclic mechanical movement). Cyclic mechanical movement is known to strongly affect the mechanical properties of viscoelastic sealants. Further studies are needed to better document the effect of mechanical movement on the ageing of building sealants. Some recent research has been done using sophisticated chemical analysis techniques. By combining mechanical and chemical testing, the ageing mechanisms of sealants may be more easily understood. In addition, the study of dynamic mechanical properties, such as the complex modulus of elasticity and the phase angle, rather than static mechanical properties, such as tensile strength and compression set, may provide direct information with regard to the capability of sealants to perform over certain service temperature ranges. 6 References 1. 2.

3. 4. 5.

6. 7.

8. 9.

Lower, L.D. (1988) Chemical influences on sealant durability, Durability, The Adhesive and Sealant Council, Inc., Arlington, U.S.A., pp. 39–49. Karpati, K.K. (1984) Investigation of factors influencing outdoor performance of two-part polysulfide sealants, Journal of Coatings Technology, Vol. 56, No. 719, pp. 57–60. Wolf, A.T. (1997) A review of the work of the RILEM Committee TC139-DBS: Durability of Building Sealants, Materials and Structures, Vol. 31, pp. 149–152. Klosowski, J.M. (1989) Sealants in Construction, Marcel Dekker Inc., New York. Lowe, G.B., Lee, T.C.P., Comyn, J. and Huddersman, K. (1994) Water durability of adhesive bonds between glass and polysulfide sealants, International Journal of Adhesion and Adhesives, Vol. 14, No. 2, pp. 85–92. Evans, R.M. (1993) Polyurethane Sealants: Technology and Applications, Technomic Publishing, Lancaster, Pennsylvania, USA. Timberlake, J.F. (1990) Silicon modified polyethers for moisture cured sealants, in: Building Sealants: Materials, Properties, and Performance, (ed. T.O’Connor), ASTM STP 1069, American Society for Testing and Materials, Philadelphia, pp. 271–281. Mansfield, C.B. (1990) Tests for the water resistance of construction sealants, Construction and Building Materials, Vol. 4, No. 1, pp. 37–42. Oba, K. and Roller, A. (1995) Characterisation of polymer-modified bituminous roofing membranes using thermal analysis, Materials and Structures, Vol. 28, pp. 596–603.


10.

11.

12.

13.

14.

15.

16. 17.

18.

19.

20.

21.

22.

23. 24.

Paroli, R.M., Delgado, A.H. and Cole, K.C. (1994) Applications of thermogravimetry and PAS–FTIR in the characterisation of silicone sealants, Canadian Journal of Applied Spectroscopy, Vol. 39, No. 1, pp. 7–14. Paroli, R.M. and Delgado, A.H. (1994) Applications of thermogravimetry-Fourier transform IR spectroscopy in the characterisation of weathered sealants, American Chemical Society Symposium Series, No. 581, pp. 129–148. Paroli, R.M., Cole, K.C. and Delgado, A.H. (1995) Evaluating the weatherability of polyurethane sealants, American Chemical Society Symposium Series, No. 598, pp. 117–136. Oba, K. and Björk, F. (1993) Dynamic mechanical properties of single-ply roof coverings for low-slope roofs and the influence of water, Polymer Testing, Vol. 12, pp. 35–56. Hanhela, P.J., Huang, R.H.E., Paul, D.B. and Symes, T.E.F. (1986) Water immersion of polysulfide sealants. II. An interpretation of the influence of curing systems on water resistance, Journal of Applied Polymer Science, Vol. 32, pp. 5415–5430. Hatsuo, I. (1996) Attenuated total reflection Fourier transform infrared spectroscopy (ATR–FTIR): a useful tool for the molecular analysis of adhesives, sealants, and coatings, Journal of Adhesives and Sealants Council, Vol. 1, pp. 521–533. Billmeyer, F.W. Jr. (1984) Textbook of Polymer Science, Third Edition, John Wiley & Sons, New York. Hugener, M. and Hean, S. (1994) Comparison of short-term ageing methods for joint sealants, in: Proceedings of the International RILEM Symposium on Durability of Building Sealants, (eds. J.Beech and A.T.Wolf), E & FN Spon, London, pp. 37–47. Tock, R.W., Dinivahi, M.V.R.N. and Chew, C.H. (1988) Viscoelastic properties of structural silicone rubber sealants, Advances in Polymer Technology, Vol. 8, No. 3, pp. 317–324. Björk, F. (1996) Methods to study degradation of rubbers in building applications, in: Durability of Building Materials and Components 7, Vol. 1, (ed. C. Sjöstrom), E & FN Spon, London, pp. 703–712. Keshavaraj, R. and Tock, R.W. (1994) Changes in crosslink density of structural silicone sealants due to ozone and moisture, Polymer-Plastics Technology and Engineering, Vol. 33, No. 4, pp. 397–417. Lee, T., Rees, T. and Wilford, A. (1992) Polysulfide sealants in water-retaining structures, in: Science Technology of Building Seals, Sealants, Glazing and Waterproofing, ASTM STP 1168, (ed. C.J.Parise), American Society for Testing and Materials, Philadelphia, pp. 47–56. Beech, J. and Mansfield, C. (1990) The water resistance of sealants for construction, in: Building Sealants: Materials, Properties and Performance, ASTM STP 1069, (ed. T.O’Connor), American Society for Testing and Materials, Philadelphia, pp. 209–220. Chew, M.Y.L. and Yi, L.D. (1997) Elastic recovery of sealants, Building and Environment, Vol. 32, No. 3, pp. 187–193. Aubrey, D.W. and Beech, J.C. (1985) The performance of joint sealants between porous surfaces in wet conditions, in: Adhesives, Sealants and Encapsulants (ASE)


25.

26.

27. 28.

29.

30.

31. 32.

1985 Conference Proceedings: Second Volume, Network Events, Buckingham, England, pp. 360–366. Myers, J.C. and Nelson, P.E. (1992) The impact of curing conditions on the appearance of sealant before and after artificial weathering, in: Science and Technology of Building Seals, Sealants, Glazing and Waterproofing, ASTM STP 1168, (ed. C.J.Parise), American Society for Testing and Materials, Philadelphia, pp. 9–21. Jacob, L.S. and Johanson, B. (1990) Structural glazing from a curtain wall designer’s perspective, in: Building Sealants: Materials, Properties, and Performance, ASTM STP 1069, (ed. T.O’Connor), American Society for Testing and Materials, Philadelphia, pp. 79–92. Tock, R.W. (1990) Temperature and moisture effects on the engineering properties of silicone sealants, ibid., pp. 167–173. Tock, R.W. and Keshavaraj, R. (1996) Physical-chemical changes in silicone elastomers used for building sealants, in: Science and Technology of Building Seals, Sealants, Glazing and Waterproofing: Fifth Volume, ASTM STP 1271, (ed. M.A.Lacasse), American Society for Testing and Materials, Philadelphia, pp. 113–125. Stoegbauer, H. and Wolf, A.T. (1990) The influence of heat ageing on one-part construction silicone sealants, in: Building Sealants: Materials, Properties and Performance, ASTM STP 1069, (ed. T.O’Connor), American Society for Testing and Materials, Philadelphia, pp. 193–208. Lerchenthal, H. and Rosenthal, I. (1972) Changes in some characteristics of polysulphide-based joint sealants in accelerated ageing tests, Materials and Structures, Vol. 5, No. 29, pp. 323–330. Sandberg, B. (1991) Comparisons of silicone and urethane sealant durabilities, Journal of Materials in Civil Engineering, Vol. 3, No. 4, pp. 278–291. Keshavaraj, R. and Tock, R.W. (1994) Oxidative effects of ozone on the ageing of structural silicone elastomers, Advances in Polymer Technology, Vol. 13, No. 2, pp. 149–156.

RECENT TRENDS IN SEALANT ADHESION TESTS An update on sealant adhesion testing N.E.SHEPHARD Sealants Science and Technology, Dow Corning Corporation, Midland, USA

Abstract A sealant is a special type of adhesive which is used in joints undergoing substantial dimensional changes. The sealant accommodates the joint movements and prevents moisture and dirt from entering the joint. In order to fulfil its function, the sealant needs to adhere to the substrate. Wherever two different materials come in contact, an interphase is formed. The strength of this interphase plays a significant role in the overall performance of the joint. Typical sealant adhesion tests require cohesive failure in the sealant as the most important indicator of a strong bond. However, this evaluation criterion only ensures that, under certain test conditions, the cohesive strength of the sealant is weaker than its adhesive strength. It does not evaluate the actual bond strength. On the other hand, a careful study of the interfacial failure can be very useful for understanding the durability of the sealant/substrate bond. This paper discusses some recent adhesion test methods which attempt to measure the strength of the interphase formed in a sealant joint. Keywords: Adhesion, durability, fatigue, fracture, interphase, sealant, viscoelasticity. 1 Introduction Sealants are designed for use in joints which undergo large dimensional changes during use [1]. Additionally, sealants are expected to function for many years. Predicting the performance of the sealant joint is critical to the sealant user. The

RECENT TRENDS IN SEALANT ADHESION TESTS 181

numerous reviews on this topic are a testimony to the importance and difficulty of this task [2–6]. Many test methods have been developed to measure the adhesion of the sealant to the various substrates and, in many requirement standards, great emphasis has been placed on the failure mode of the sealant to the substrate. Cohesive failure in the sealant has long been considered a desirable feature. Cohesive failure occurs when the fracture energy of the sealant is less than the fracture energy of the interphase between the sealant and the substrate. Adhesion tests are almost always used to compare one sealant formulation to another. However, if both sealants fail cohesively during the test, then a simple statement of failure mode will not distinguish between the relative durability of the adhesion of the two sealants. However, recently suggestions have been made to modify the adhesion test methods in order to increase the likelihood of adhesive failure. This is usually accomplished by applying a stress which is less than the cohesive strength of the sealant. This criterion ensures that cohesive failure is unlikely to occur. Interfacial failure is then monitored during or after exposure to accelerated weathering conditions. 2 Peel test methods The most common peel test is the 180° peel test conducted at constant peeling speed. The force per unit peel width and the failure mode are recorded. This particular peel test is very difficult to interpret because the failure mode and peel strength are strongly influenced by the viscoelastic properties of the sealant. Because of this dependency, both the failure mode and the peel strength can be substantially altered by changing the experimental conditions, especially the peeling speed and the sealant bead thickness [7]. Obviously, the sealant modulus affects the peel force and, therefore, the failure mode during peel testing [8]. Figure 1 demonstrates the relationship between sealant modulus, peeling speed, sealant bead thickness and failure mode. The graph shows the cohesive failure for 180° peel tests using a room temperature curing silicone sealant adhered to glass and aluminium. No adhesion promoters were used in this sealant formulation [8]. Slower peeling speed and thinner sealant beads are suggested for increased sensitivity to adhesive failure. Since the failure mode and peel force are strongly governed by the peeling speed, it is necessary to characterise the adhesion of a sealant using peeling speeds which are expected in the field. Sealants undergo cyclic strain caused by daily and seasonal changes in temperature. These movements occur at a rather slow speed; the resulting rate of strain may be 10–6 m/sec or slower. A device for measuring crack speeds at the sealant/substrate interphase at this rate under constant load peel conditions has been demonstrated [9]. In this device, gravity is used to supply a constant peeling load to a peeling specimen


Fig. 1. Region of cohesive failure for 180° peel tests using silicone sealant cured to glass and aluminium.

mounted at a 45° angle to the peeling force using a simple lever and fulcrum. The crack speed is continuously measured using a linearly variable displacement transducer and a computer for data acquisition. The experiment then consists of conducting many measurements at different applied load and plotting the resulting crack speed versus fracture energy graph on a logarithmic scale. The improvement over the standard 180° peel test is readily apparent from Figure 2. The constant load peel test can be used to easily differentiate the quality of the adhesion between the two different sealants, while the traditional 180° peel test indicates both sealants have good adhesion. In the constant load test, one of the two sealants failed adhesively with very low peel strength, while both sealants failed cohesively in the traditional 180° peel test. This suggests that the standard 180° peel test has a tendency to yield passing results, even when the sealant has poor interfacial strength. The constant load peel test can also be used to determine a rate equation for crack speed as a function of peel force, relative humidity and temperature by exposing the peel specimen to heat and moisture during the test [10]. By measuring the crack speed as a function of applied peeling force and at constant temperature and humidity, series of curves are obtained for various temperatures and relative humidities as shown in Figure 3. A power law rate equation can then be deduced using the method of reduced variables (commonly called timetemperature superposition). Using this power rate equation, a master curve is


Fig. 2. Logarithmic plot of crack speed versus the applied fracture energy for two different sealants (x: sealant A, o: sealant B; sealant A exhibits cohesive failure, sealant B adhesive failure in the constant load test. The triangles and squares in the upper right corner represent the results of the 180° peel test, in which both sealants failed cohesively [9]).

constructed by shifting the curves. Finally, from the resulting shift factor plots, an equation for crack speed as a function of peeling force can be obtained: (1) where a is the crack speed, arh and aT are the shift factors for relative humidity and temperature, G is the applied fracture energy. The temperature shift factor can be determined by an Arrhenius approach. The relative humidity shift factor is obtained by empirical curve fitting of the shift factor plot. Such equations are very useful for developing predictive models for sealant durability. Furthermore, because of the time-temperature superposition, all experimental data can be obtained over a very reasonable period of time, such as days or weeks. Once shifted and converted to a master curve, the adhesive strength can be estimated over twenty decades of time, while the collection of real time data over this range of time is clearly impractical. 3 Tensile adhesion test methods Tensile adhesion tests are often conducted by preparing a tensile adhesion specimen in the usual fashion. The stress and strain to break of the specimen is measured at a constant pull speed. Again, the failure mode is also reported.


Fig. 3. Doubly shifted master curves for fracture energy as a function of crack speed [10].

Cohesive failure is desired in order to demonstrate good interfacial strength between the sealant and the substrate. This is a classic example of a continuum mechanics approach to adhesion. 3.1 Continuum mechanics approach and the fracture mechanics approach Continuum mechanics methods assume that the joint is free of flaws or imperfections [11]. The average far field stress in the joint is used to determine the criteria for joint failure. In the case of the tensile adhesion joint, the far field stress is the applied load divided by the cross sectional area of the joint. Stress concentrations due to intrinsic macroscopic defects are ignored or averaged. The failure is assumed to be catastrophic. The design stress is then determined by measuring the average stress needed to break the joint and multiplying this value by a safety factor. It is well known that average far field stresses can be greatly magnified near the locus of a defect in the joint. Such zones of increased stress are often referred to as stress concentrations. Defects can be obvious, such as air pockets, cracks or


foreign particle inclusions. Stress concentrations occur whenever the mechanical properties of the joint are significantly changed. Rather than evaluating average stresses, fracture mechanics studies the energy needed to propagate a known flaw through the solid. Typically, a flaw is deliberately placed in the joint so it can be studied. The fracture mechanics of rubbery materials has been recently reviewed [12]. 3.2 Recent applications of adhesion methods using tensile adhesion joints A simple yet effective approach to improve the utility of the tensile adhesion joint has been to conduct the test using a fixed displacement (constant strain) and monitoring the time to failure under various exposure conditions. Two issues exist for constant strain tensile adhesion tests. First, the applied forces are directly related to the modulus of the sealant. Therefore, low modulus sealants perform well in these tests, even if their interfacial strength is low. Second, sealants which undergo chemical stress relaxation will perform unusually well, because the applied forces will decay to zero in a short period of time. In order to study the durability of sealant joints, the constant strain tensile adhesion test has been combined with outdoor exposure and cyclic movement [13]. This method is very effective at predicting joint failure because stress relaxation and compression set are allowed to occur naturally. Recently, a tensile adhesion test has been conducted that studies tensile adhesion joints under constant load while immersed in hot water 13]. This method improves the resolution of the tensile adhesion test by maintaining a constant load on the joint. In all of this work, the fracture energy of the sealant was not reported. Only average tensile strength values are reported because the testing was conducted using continuum mechanics methods. In another study, the cohesive fracture energy of silicone sealant was measured in a tensile adhesion joint [15]. This method is similar to the constant displacement tensile adhesion method, but was modified to measure crack speed and fracture energy. The fracture energy is calculated by measuring the strain energy stored in the joint at a given strain. In the study, a cut was placed in the sealant near the centre point of one edge to initiate crack growth. The speed of the crack growth was monitored visually. This test can be made more sensitive to the strength of the interphase by placing the pre-crack along the edge of the sealant and the interface between the sealant and the substrate. Using this modified test, the interfacial crack growth rate versus applied fracture energy was measured and compared to peel test data [10]. A reasonable correlation was obtained as seen in Figure 4. Using Equation 1 and strain energy data from the tensile adhesion specimen, a hypothetical joint could be evaluated for various weathering conditions. The hypothetical joint was composed of the same sealant and aluminium substrate


Fig. 4. Fracture energy for the pure shear (o) and the peel test (x) [10].

Fig. 5. Crack length for a hypothetical expansion joint in Wittman, Arizona (solid line) and Miami, Florida (dashed line) [10],

that was used for the tensile adhesion testing. Annual weathering data for Wittman, Arizona and Miami Florida was used in conjunction with the shift factors from Equation 1. The applied fracture energy was determined from the tensile adhesion test specimen geometry and an estimation of the size and thermal expansion of the joint as a function of temperature changes. The crack length versus time for a hypothetical tensile adhesion joint was calculated for the two different cities and can be found in Figure 5. From this crude analysis, several trends were noted. Crack growth mainly occurs in the winter months when the joint is under tensile forces. The larger crack growth seen in Wittman could be attributed to the colder winter months. Furthermore, the Miami crack growth was a smoother function because the relative humidity is always high in Miami. Conversely, Wittman has very low humidity and crack growth only occurs on cold wet days.


Fig. 6. Cycles to failure for a silicone sealant as a function of strain amplitude [16].

A recent example of fracture mechanics applied to cyclic stress is a further enhancement of the tensile adhesion joint test. The cohesive fracture energy of silicone sealant was measured as a function of fatigue cycles at a fixed frequency and amplitude [16]. Life predictions where calculated from this data as seen in Figure 6. From this data, it is easy to see that the sealant life will be strongly governed by the maximum joint movement that occurs. However, this fatigue data was only obtained for cohesive crack growth. Future work should consider starting the crack at the locus of the interface between the sealant and the substrate. 4 Summary and suggestions for future work The culmination of the recent adhesion work may be a test which uses the methods of interfacial fracture mechanics on tensile adhesion joints subjected to cyclic stress along with shift factor plots to account for weathering factors. Such a test methodology would result in the generation of a single equation for cycles to failure which would be a function of the applied stresses and combined weathering affects. In conclusion, sealant adhesion test methods are being modified to mimic strain rates and stresses seen in actual use. Consequently, the test methods have become much more sensitive to the sealant/substrate interface. The trend is to use applied loads and strains which are less than the forces need to initiate cohesive failure. Adhesive failure is then monitored as a function of artificial weathering. Additionally, fracture mechanics methods are becoming more popular. A combination of fracture mechanics to quantify interfacial crack speed as a function of cyclic stress along with artificial weathering will soon be common. Such methods of measurement should provide useful rate equations


which can be used in computer aided design of sealant joints for superior durability. 5 References 1. 2.

3.

4. 5.

6.

7.

8.

9.

10. 11. 12. 13.

Klosowski, J.M. (1989) Sealants in Construction, Marcel Dekker Inc., New York and Basel. Gutowski, V.S., Lalas, P. and Cerra, A.P. (1996) Structural silicone in curtain walls, in: Science and Technology of Building Seals, Sealants, Glazing and Waterproofing: Fifth Volume, ASTM STP 1271, (ed. M.Lacasse), American Society for Testing and Materials, Philadelphia, pp. 97–112. Beech, J.C. (1996) The contribution of research to standardisation, in: Durability of Building Materials, Proceedings of the International RILEM Symposium, Vol. 28, (ed. J.C.Beech and A.T.Wolf), E & FN Spon, London, pp. 123–134. Wolf, A.T. (1996) Ageing resistance of building and construction sealants, Part 1, ibid., pp. 63–89. Lacasse, M.A. (1994) Advances in test methods to assess the long-term performance of sealants, in: Technology of Building Seals, Sealants, Glazing and Water-proofing: Third Volume, ASTM STP 1254, (ed. J.Myers), American Society for Testing and Materials, Philadelphia, pp. 5–20. Wolf, A.T. (1990) Studies of the ageing behaviour of gun-grade building joint sealants—the “state-of-the-art”, Polymer Degradation and Stability, Vol. 23, pp. 135–163. Gutowski, V.S., Russell, L. and Cerra, A. (1992) New tests for adhesion of silicone sealants, in: Science and Technology of Building Seals, Sealants, Glazing and Waterproofing: Second Volume, ASTM STP 1200, (ed. J.Klosowski), American Society for Testing and Materials, Philadelphia, pp. 87–104. Shephard, N.E. and Wightman, J.P. (1995) An analysis of the 180° peel test for measuring sealant adhesion, in: Science and Technology of Building Seals, Seal ants, Glazing and Waterproofing: Fifth Volume, ASTM STP 1271, (ed. M. Lacasse), American Society for Testing and Materials, Philadelphia, pp. 226–238. Shephard, N.E. and Wightman, J.P. (1998) A simple device for measuring adhesive failure to sealant joints, in: Science and Technology of Building Seals, Sealants, Glazing and Waterproofing: Seventh Volume, ASTM STP 1334, (ed. J. Klosowski), American Society for Testing and Materials, Philadelphia (in print). Shephard, N.E. (1995) Measuring and predicting sealant adhesion, Dissertation Abstracts International Series B, Vol. 56, No. 5, p. 2814. Anderson, G.P., Bennett, S.J. and DeVries, K.L. (1977) Analysis and Testing of Adhesive Bonds, Academic Press, New York. Lake, G.J. (1995) Fatigue and fracture of elastomers, Rubber Chemistry and Technology, Vol. 68, No. 3, pp. 435–442. Lacasse, M.A., Bryce, J.E. and Margeson, J.C. (1995) Evaluation of cyclic fatigue as a means of assessing the performance of construction joint sealants: silicone sealants, in: Science and Technology of Building Seals, Sealants, Glazing and Waterproofing: Fourth Volume, ASTM STP 1243, (ed. D.Nicastro), American Society for Testing and Materials, Philadelphia, pp. 49–64.


14.

15.

16.

Cerra, A.P. and Gutowski, V.S. (1996) Performance-based adhesion testing of structural sealants, in: Science and Technology of Building Seals, Sealants, Glazing and Waterproofing: Fifth Volume, ASTM STP 1271, (ed. M.Lacasse), American Society for Testing and Materials, Philadelphia, pp. 209–225. Mazich, K.A., Samus, M.A., Smith, C.A. and Rossi, G. (1991) Threshold fracture of lightly cross linked networks, Macromolecules A, Vol. 24, No. 10, pp. 2766–2769. Lacasse, M.A., Margeson, J.C. and Dick, B.A. (1996) Static and dynamic cut growth fatigue characteristics of silicone-based elastomeric sealants, in: Durability of Building Materials, Proceedings of the International RILEM Symposium, Vol. 28, (ed. J.C.Beech and A.T.Wolf), E & FN Spon, London, pp. 1–16.

EVALUATION OF SILICONE SEALANTS EXPOSED TO NATURAL AND ACCELERATED WEATHERING A comparative evaluation of weathering techniques S.SUGIYAMA Dow Corning Toray Silicone Co., Ltd., Chiba, Japan

Abstract The tensile adhesion properties of two silicone sealants were determined after exposure of static sealant specimens to three accelerated weathering regimes and to outdoor weathering. The accelerated weathering was carried out for a duration of up to 15,000 hours in commercial weathering machines based on carbon flame, xenon lamp and fluorescent lamp (QUV) light sources. The sealants were exposed to outdoor weathering for up to three years at Ichihara, Chiba, Japan. Very little change in modulus was found for specimens exposed to outdoor weathering. Rather close correlation between the results obtained in the various accelerated weathering methods was observed. The results obtained in accelerated weathering and outdoor exposure do not correlate as well, however, the trends in property changes are well reproduced. Keywords: Accelerated weathering, carbon flame, outdoor exposure, QUV, silicone sealant, xenon lamp. 1 Introduction Building materials, such as sealants, are expected to have long service lives in order to minimise the life cycle cost of a building. A reliable estimation of the life expectancy of a sealant is therefore desirable and the use of accelerated weathering methods for this purpose has become widespread. Certain accelerated weathering methods are described in national or international standards. However, our understanding of the degradation mechanisms occurring in

EVALUATION OF SILICONE SEALANTS EXPOSED 191

sealants is not yet well enough developed to allow for a good prediction of the actual in-service performance of sealants based on their behaviour in accelerated weathering tests. A number of environmental factors influence the durability of sealants. Especially important are the effects of ultraviolet (UV) and visible light as well as high temperature, water, oxygen and mechanical cycling. In actual service, these factors occur to varying degrees, depending on the type of joint and the exposure conditions. Some of the environmental factors act synergistically, making the determination of the degradation mechanism particularly difficult. Weathering machines based on a carbon arc flame light source are the most common type of accelerated ageing equipment found in Japan and have been frequently used to study the durability of sealants. However, not much emphasis has been placed in past studies on the correlation of the results obtained in carbon arc equipment with outdoor exposure or other accelerated weathering machines. The purpose of this study therefore was to investigate the difference in the degradation behaviour of two silicone sealants under three different types of accelerated weathering and outdoor exposure. 2 Experimental 2.1 Test materials Two silicone sealants were evaluated in this study. Sealant Si-A was a chalk filled, one-part, alkoxy cure, medium to high modulus silicone sealant. Sealant Si-B was a chalk filled, two-part, aminoxy cure, low modulus silicone sealant. Both sealants are commonly used in curtain wall glazing applications in Japan. 2.2 Preparation of test specimen Tensile adhesion test specimens were prepared according to JIS A 5758–1992 [1] (Figure 1) using float glass as substrate. Sealant Si-A was cured for two weeks at 20°C and 55% R.H. followed by two weeks at 30°C (at ambient humidity). Sealant Si-B was cured one week at 20°C and 55% R.H. followed by one week at 50°C (at ambient humidity). Triplicate specimens were prepared for each weathering conditions. The specimens were exposed in static conditions (no movement or pre-stress) such that the artificial or natural light was shining through one of the glass supports irradiating the sealant/glass interface.


Fig. 1. Tensile adhesion specimen according to JIS A 5758 standard.

2.3 Outdoor weathering The outdoor weathering was carried out according to JIS A 1410 [2]. The specimens were placed at an angle of 45° facing South on the roof of a four story building in Ichihara city, Chiba, Japan. 2.4 Accelerated weathering The specimens were exposed to carbon flame and xenon lamp type accelerated weathering according to JIS A 1415 test standard [3] and to fluorescent UV lamp type accelerated weathering according to ASTM G–53 test practice [4]. In the carbon flame and xenon lamp type equipment, the black panel temperature was kept at 63°C and the specimens were sprayed with water for 18 minutes every 2 hours while the specimens were being irradiated. A cycle consisting of four hours of UV irradiation at 60°C and four hours of dark condensation period at 40°C was used in the QUV fluorescent type equipment. All above mentioned light sources are defined in the JIS K 7350 standard [5]. 2.5 Tensile adhesion test Tensile adhesion tests were carried out on triplicate specimens according to JIS A 5758 standard with an extension rate of 50 mm/min.


Table 1. Tensile adhesion properties of sealant Si-A after outdoor weathering

3 Result and discussion The results are summarised in Tables 1–8. The values represent the average of test data obtained on triplicate specimens. 3.1 Result of outdoor weathering The tensile adhesion properties of the specimens exposed to outdoor weathering are shown in Tables 1 and 2 for sealants Si-A and Si-B. Note that M50 indicates the tensile secant modulus of the sealant at 50% elongation, Tmax and Emax are its tensile strength and elongation at break, respectively, while CF and AF indicate cohesive and adhesive failure mode, respectively. 3.2 Result of accelerated weathering The tensile adhesion properties of sealants Si-A and Si-B after accelerated weathering are shown in Tables 3–5 and Tables 6–8, respectively. Table 2. Tensile adhesion properties of sealant Si-B after outdoor weathering Exposure time (h)

M50 (MPa)

Tmax (MPa)

Emax (%)

Failure mode CF%

0 4380 8760 26280

0.216 0.177 0.196 0.186

1.079 0.814 0.824 0.804

980 1100 1056 1110

100 100 100 100

AF% 0 0 0 0


Table 3. Tensile adhesion properties of sealant Si-A after exposure to Xenon lamp type accelerated weathering Irradiation time (h)

0 1000 3000 5000 8000 10000 15000

M50 (MPa)

0.304 0.275 0.284 0.275 0.284 0.284 0.304

Tmax (MPa)

1.158 0.883 0.804 0.726 0.687 0.736 0.657

Emax (%)

347 223 193 196 172 187 158

Failure mode AF%

CF%

100 100 100 100 100 100 100

0 0 0 0 0 0 0

Table 4. Tensile adhesion properties of sealant Si-A after exposure to carbon arc flame type accelerated weathering Irradiation time (h)

0 1000 3000 8000 10000 15000

M50 (MPa)

0.304 0.275 0.284 0.275 0.265 0.265

Tmax (MPa)

1.158 0.804 0.824 0.697 0.706 0.706

Emax (%)

347 198 189 186 182 179

Failure mode CF%

AF%

100 100 100 100 100 100

0 0 0 0 0 0

Table 5. Tensile adhesion properties of sealant Si-A after exposure to QUV fluorescent lamp type accelerated weathering Irradiation time (h)

M50 (MPa)

Tmax (MPa)

Emax (%)

Failure mode CF%

0 1000 3000 5000 8000 10000 15000

0.304 0.275 0.265 0.275 0.275 0.255 0.255

1.158 0.893 0.804 0.824 0.706 0.726 0.736

347 229 192 192 167 207 215

100 100 100 100 100 100 100

AF% 0 0 0 0 0 0 0


Table 6. Tensile adhesion properties of sealant Si-B after exposure to Xenon lamp type accelerated weathering Irradiation time (h)

M50

Tmax

(MPa)

(MPa)

(%)

0 1000 3000 5000 8000 10000 15000

0.216 0.186 0.157 0.157 0.147 0.157 0.157

1.079 0.883 0.746 0.716 0.667 0.706 0.647

Emax

Failure mode CF%

980 1139 1144 1182 1124 1141 1122

AF%

100 100 100 100 100 100 100

0 0 0 0 0 0 0

Table 7. Tensile adhesion properties of sealant Si-B after exposure to carbon arc flame type accelerated weathering Irradiation time (h)

M50

Tmax

(MPa)

(MPa)

(%)

0 1000 3000 5000 8000 10000 15000

0.216 0.186 0.157 0.167 0.147 0.147 0.157

1.079 0.863 0.755 0.755 0.736 0.647 0.716

Emax

Failure mode CF%

980 1104 1163 1124 1161 1146 1139

AF%

100 100 100 100 100 100 100

0 0 0 0 0 0 0

Table 8. Tensile adhesion properties of sealant Si-B after exposure to QUV fluorescent lamp type accelerated weathering [rradiation time (h)

M50

Tmax

(MPa)

(MPa)

(%)

0 1000 3000 5000 8000 10000 15000

0.216 0.137 0.147 0.137 0.137 0.137 0.137

1.079 0.795 0.628 0.736 0.706 0.569 0.608

Emax

Failure mode CF%

980 1252 1224 1228 1208 1152 1204

100 100 100 100 100 100 100

AF% 0 0 0 0 0 0 0


Fig. 2. Changes in tensile properties of sealant Si-A during outdoor weathering.

Fig. 3. Change in M50 of sealant Si-A during accelerated weathering.

3.3 Comparison of outdoor and accelerated weathering The changes in M50, Tmax and Emax for sealant Si-A after outdoor and accelerated weathering are shown in Figures 2–5. As can be seen from Figure 2, outdoor weathering causes little change in M50, while Tmax and Emax were found to decrease. Accelerated weathering causes some decrease in M50 (Figure 3) and larger decreases in Tmax (Figure 4) and Emax (Figure 5). For all weathering types, the majority of the change in tensile properties occurs at the initial stage of weathering (within the first 5,000 hours of exposure). The changes in M50, Tmax and Emax for sealant Si-B after outdoor and accelerated weathering are shown in Figures 6–9. As can be seen from Figure 6, outdoor weathering causes some decrease in M50 and Tmax, while Emax was found to increase somewhat. Accelerated weathering makes these changes more pronounced, however, the same trends are observed. As for sealant Si-A, most of the change in tensile properties occurred within the first 5,000 hours of exposure.


Fig. 4. Change in Tmax of sealant Si-A during accelerated weathering.

Fig. 5. Change in Emax of sealant Si-A during accelerated weathering.

The difference in the ageing behaviour of sealants Si-A and Si-B can probably be attributed to differences in the cure systems employed in the two sealants, specifically to the residual activity of the catalysts. 3.4 Correlation of the various weathering methods The most striking finding of this study is how closely the results correlate between the various accelerated weathering methods. Figure 10 shows, as an example, the correlation between the tensile strength values found in carbon arc flame and xenon lamp type accelerated weathering. The results obtained in accelerated weathering and outdoor exposure do not correlate as well as between the various accelerated weathering methods, however, the trends in property changes are well reproduced. This suggests that light, both visible and ultraviolet, is not the primary ageing factor for silicone sealants. The rather small differences observed between the results of the various accelerated weathering methods are more likely caused by differences in the heat


Fig. 6. Changes in tensile properties of sealant Si-B during outdoor weathering.

Fig. 7. Change in M50 of sealant Si-B during accelerated weathering.

Fig. 8. Change in Tmax of sealant Si-B during accelerated weathering.

and water exposure (for instance, water spraying versus water condensation). It should be noted that for organic sealants, which are much more sensitive to light exposure, the rather good correlation between the various accelerated weathering methods may not hold true. On the contrary, since some organic polymer


Fig. 9. Change in Emax of sealant Si-B during accelerated weathering.

Fig. 10. Correlation of Tmax changes observed for sealant Si-A between Xenon lamp and carbon arc flame.

systems are known to be extra sensitive in specific wavelength regions, it is very likely that different light sources will provide different weathering results. 4 Summary The tensile adhesion properties of two silicone sealants were determined after exposure of static sealant specimens to three accelerated weathering regimes and to outdoor weathering. The accelerated weathering was carried out for a duration of up to 15,000 hours in commercial weathering machines based on carbon flame, xenon lamp and fluorescent lamp (QUV) light sources. The sealants were exposed to outdoor weathering for up to three years at Ichihara, Chiba, Japan. Very little change in modulus was found for specimens exposed to outdoor weathering.


Rather close correlation between the results obtained in the various accelerated weathering methods was observed. The results obtained in accelerated weathering and outdoor exposure do not correlate as well, however, the trends in property changes are well reproduced. This suggests that light, both visible and ultraviolet, is not the primary ageing factor for silicone sealants. The rather small differences observed between the results of the various accelerated weathering methods are more likely caused by differences in the heat and water exposure (for instance, water spraying versus water condensation). The difference in the ageing behaviour of sealants Si-A and Si-B can probably be attributed to differences in the cure systems employed in the two sealants, specifically to the residual activity of the catalysts. 5 References 1.

2. 3. 4.

5.

Japanese Industrial Standards Committee (1992) Sealing Compounds for Sealing and Glazing in Building. JISC, Standards Department, Ministry of International Trade and Industry, Tokyo. JIS A 5758. Japanese Industrial Standards Committee (1968) Recommended Practice for Outdoor Exposure of Plastics Building Materials, ibid., JIS A 1410. Japanese Industrial Standards Committee (1994) Recommended Practice for Accelerated Artificial Exposure of Plastics Building Materials, ibid., JIS A 1415. American Society for Testing and Materials (1996) Standard Practice for Operating Light and Water Exposure Apparatus (fluorescent UV-condensation type) for Exposure of Non-metallic Materials, ASTM, Philadelphia, ASTM G-53. Japanese Industrial Standards Committee (1996) Plastics—Methods of Exposure to Laboratory Light Sources. JISC, Standards Department, Ministry of International Trade and Industry, Tokyo. JIS K 7350 (adopted from ISO 4892).

TEST METHODS ASSESSING THE EFFECT OF SEALANTS ON THE AESTHETIC APPEARANCE OF BUILDINGS Efforts towards the development of international standards A.T.WOLF Dow Corning S.A., Seneffe, Belgium

Abstract The appearance of any building changes with time as a result of complex processes generally summarised under the term “ageing”. While sealants can contribute to these processes, many aesthetic issues will appear on a building facade over time with or without the presence of a sealant. A work group (WG10) has been charted within the International Standardisation Committee TC59/SC8 (Building Construction, Jointing Products) to assess the need for developing test standards on aesthetic issues caused by sealants. Two work items are being pursued for standardisation: “Staining of Porous Substrates” and “Determination of Surface Changes Following Exposure to Artificial or Natural Weathering”. The paper discusses key aesthetic issues occurring with sealants and the proposed test methods. Keywords: Chalking, crazing, fluid migration, sealant, staining, surface changes. 1 Introduction The appearance of any building changes with time, this is the result of complex processes generally summarised under the term “ageing”. A building’s aesthetics may be either enhanced by ageing, such as by the development of a patina on a copper roof, or degraded, such as by algal growth on areas affected by moisture. While sealants can contribute to these processes, many aesthetic issues will appear on a building facade over time with or without the presence of a sealant.


In 1995, the International Standardisation Committee TC59/SC8 (Building Construction, Jointing Products) formed a Work Group (WG10) to assess the need for developing test standards on aesthetic issues caused by sealants. The Work Group started by collecting information on the various types of aesthetic issues caused by sealants. Nine aesthetic issues were identified and prioritised in their sequence of importance. The Work Group then critically evaluated the various national test standards used for assessing aesthetic issues with sealants and reviewed the literature on the topic. As a result of these discussions, the author prepared a state-of-the-art report [1], proposing definitions of the various aesthetic issues observed with weatherproofing sealants, documenting our understanding of the mechanisms causing aesthetic issues, and proposing potential test methods for the qualitative or quantitative assessment of the aesthetic performance of sealants. Within three meetings of the Work Group, it became apparent that sufficient test experience and scientific understanding exist to pursue the following two work items for the development as an International Standard: “Staining of Porous Substrates” and “Determination of Surface Changes Following Exposure to Artificial or Natural Weathering”. The Work Group is currently evaluating various proposed test methods for these two work items and hopes to achieve a Committee Draft (ISO-CD) by 1999/2000 and a Draft International Standard (ISO-DIS) within further two years. This paper attempts to convey the basic test concepts that are being considered within ISO Work Group 10 along with some of the key problems and open questions that remain. 2 Definition of terminology Aesthetic issues raised by sealants may be classified into two groups: those that affect the sealant itself and those that affect the adjacent building substrates. Aesthetic issues that fall within the first category are blooming, dirt pick-up, microbial growth, chalking, surface crazing, change of colour and change of gloss. Aesthetic issues which affect the adjacent building substrates are fluid migration and fluid streaking. The following definitions of the above aesthetic issues are being proposed [1]: • blooming: accumulation of formulation components on a sealant surface; • dirt pick-up: soiling caused by foreign material that is deposited on or embedded into a sealant surface; • microbial growth: visible accumulation and colonisation of micro-organisms on a sealant surface; • chalking: whitening or formation of a powder due to degradation and erosion of a sealant surface caused by weathering; • surface crazing: formation of cracks in a sealant surface due to weathering;

TEST METHODS ASSESSING THE EFFECT 203

• change of colour. change of colour of a sealant surface due to weathering or other environmental or external factors; • change of gloss: change of gloss of a sealant surface due to weathering or other environmental or external factors; • fluid migration: accumulation of liquid formulation components from a sealant on or in an adjacent building material, causing the development of a visible stain or aesthetically detrimental effect on the adjacent substrate; • fluid streaking: accumulation of dirt in discreet water run-off channels formed on vertical or sloped building facades as a result of the modification of the substrate surface by fluid migration from the sealant; the phenomenon is typically visible above and below horizontal joints. 3 Prioritisation of aesthetic issues for the development of test standards In its first meeting, WG 10 collected information on the various types of aesthetic issues caused with sealants and prioritised them in their sequence of importance. The result of this prioritisation is shown in Table 1. Table 1. Result of prioritisation High Priority

Medium Priority

Low Priority

Fluid Migration Dirt Pick-up Chalking Surface Crazing

Microbial Growth Fluid Streaking

Blooming Change of Colour Change of Gloss

From further discussions, it became readily apparent that dirt pick-up, while being rated a highly important aesthetic feature of the sealant, was also the most difficult to reproduce under controlled conditions. During the 1980’s, scientists in Japan had developed various methods of accelerating the dirt pick-up on sealants under laboratory conditions [2, 3]. However, in order to conduct their experiments, they had to choose an “artificial type of dirt”, which substantially differed in its chemical nature from “real dirt” (hydrophilic versus hydrophobic). Due to this experimental restriction, they obtained only a limited correlation between the artificially accelerated laboratory experiments and the performance observed on actual service joints [1]. Essentially the same problem as with dirt pick-up arises when one wishes to accelerate the occurrence of fluid streaking: the type of “artificial dirt” which needs to be involved in order to make the experiment “work” is not representative of the “real dirt” affecting service joints. Because of the difference in the chemical nature of the dirt and in the duration allowed for dirt deposition


and retention in the accelerated laboratory experiments, only a limited correlation is obtained with the performance observed in actual service joints [4]. A similar situation exists for testing the resistance of sealants to microbial growth. While it is quite simple to achieve reproducible results in laboratory experiments, the microbial resistance observed on sealants in actual service joints differs considerably from the laboratory observations, and is highly affected by the specific exposure conditions [5]. While it is highly desirable to develop internationally standardised test methods for dirt pick-up, fluid streaking and microbial growth, a more detailed scientific assessment of the current national or industrial test methods and practices needs to occur, before this task can be undertaken. The ISO TC59/SC8 Work Group 10 will continue to provide a forum for these discussions. On the other hand, the Work Group felt that two work items should be pursued for International Standardisation: The British delegation proposed that, building on the experience of the paint industry, a test method should be developed that incorporated both natural and accelerated weathering and evaluated the sealants for their resistance to chalking, surface crazing, change of colour and change of gloss. Consequently, a first draft of such a test method was developed by the British delegation. Furthermore, the American delegation suggested pursuing a test standard for the evaluation of the fluid migration (staining) potential of sealants on porous substrates. Several national test methods had been developed to this end with probably the most advanced method being the ASTM 1248 standard [6]. Consequently, a first draft of such a test method was developed by the American delegation. The two work items were submitted for voting by all ISO full member countries and received the necessary support for further development. 4 Terminology and underlying mechanism of phenomena 4.1 Surface changes following exposure to artificial or natural weathering 4.1.1 Chalking The term chalking refers to the whitening of some organic sealant surfaces with weathering. The whitening is caused by the degradation and erosion of the polymer, which leads to increased roughness or porosity of the sealant surface. The early stages of chalking are marked by a loss of surface gloss and colour.


4.1.2 Surface crazing This term refers to cracks formed on a sealant surface as a result of environmental ageing. Certain sealant types exhibit characteristic crack patterns, for instance, some polysulfide sealants exhibit parallel surface cracks perpendicular to the direction of joint movement, while some polyurethane sealants develop a pattern of cracks with ageing which is generally referred to as “mud cracking”. Surface crazing is caused by the degradation of the sealant surface induced by the shorter wavelength spectrum of the sunlight. Most surface cracks are only a few tenth of a millimetre deep, since sunlight does not penetrate deeply into sealants, which contain opaque fillers, such as calcium carbonate. However, under cyclic joint movement, surface cracks open during the expansion phase, exposing new sealant surface to sunlight. Depending on the joint movement amplitude, the exposure conditions, the type of sealant and its formulation, this mechanism can allow cracks to penetrate deeper into the sealant bulk [7]. 4.1.3 Change of colour This term is self-explanatory, however, there are several mechanisms which can cause changes in colour. A number of mechanisms are inherent in the sealant, for instance changes in colour may arise from insufficient colour fastness of one of the pigments used in the sealant formulation, from dirt pick-up or microbial growth, or from chalking. The most important external factor to cause changes in colour is the incompatibility with other building materials. 4.1.4 Change of gloss As before, this term is self-explanatory and the same mechanisms that cause changes in colour may also cause changes in gloss. 4.2 Staining of porous substrates (fluid migration) The term refers to the migration of fluids (plasticizers, unreacted or unreactive polymers, oligomers, antioxidants, or other liquid formulation additives) out of a sealant and into or onto a substrate. The phenomenon is most visible on porous or microporous substrates, where the migration of fluid(s) causes a dark band (“wet look”) around the joint. The fluid soaked band may extend from a few millimetres to several centimetres from the joint, depending on the amount and the viscosity of the material bled from the sealant.


The extent and rate of fluid migration into porous substrates depend primarily on the sealant’s formulation and secondly on its cure chemistry [8], Contrary to common belief, the chemical nature of a sealant’s polymer has little or no effect on the degree of fluid migration observed. Consequently, fluid migration has been observed for any sealant type: oleo-resinous putties, butyls, acrylics, polysulfides, polymercaptanes, polyurethanes, silicon modified polyethers, and silicones [9–14]. Since fluid migration is a diffusion controlled phenomenon, it is affected by the availability of free fluids in the sealant formulation, the “internal pressure” of the free fluids and the temperature. An increase in the ambient temperature and a compression of the sealant will accelerate the fluid migration process. These are precisely the environmental conditions a building joint sealant will be exposed to during the hot summer months. 5 Proposed test methods 5.1 Surface changes following exposure to artificial or natural weathering The proposed test standard specifies methods for the determination of surface changes (gloss, colour, chalking and cracking) on sealants after exposure to artificial or natural weathering. It also specifies a method of recording surface soiling, caused by dirt pick-up or microbial growth, after exposure to natural weathering. The specimens are prepared such that the sealant adheres to an aluminium panel. Under the artificial weathering regime, the specimens are subjected to a cyclic exposure to artificial light at elevated temperature and to water. The currently proposed test standard suggests a period of 1,000–2,000 hours of accelerated weathering comprising 500–1,000 cycles of wet and dry conditions. Exposure to natural weathering is carried out in accordance with ISO 2810 [15] on an exposure rack at an inclination angle of 45° to the horizontal with the rack facing the equator. The currently proposed test standard suggests a period of one year for the outdoor exposure. After completion of the weathering periods, changes in gloss and colour are evaluated quantitatively in accordance with ISO 2813 [16] and ISO 7224 [17], while changes in cracking and chalking are evaluated semi-quantitatively by comparing the exposed sample to a set of benchmark samples (as in ISO 4628 [18]). For the outdoor exposed samples, the degree of surface soiling may be rated in accordance with the guidance given in ISO 4628. The main limitation of this test method—as with any other method of accelerated weathering—is that valid correlations between ageing observed in natural and accelerated weathering generally cannot be obtained. Certain


relationships may be deduced, if the effect of key weathering parameters on the sealant is known [19]. However, even with the most judicious selection of weathering parameters used in the accelerated weathering, only very limited correlation between the effects of accelerated and natural weathering is to be expected, simple because our current artificial weathering technology is not capable of reproducing the complete set of weathering factors. Rather, the current weathering machines attempt to reproduce a few key factors (UV light, water and heat) simultaneously. While these factors generally suffice for reproducing the effects of weathering on the bulk properties of material, the effect of weathering on the surface properties of materials is far more complex, since more complicated synergies exist. Consider, for instance, the negative synergy between dirt pick-up and UV exposure: any dirt deposited on a sealant surface effectively shields the underlying area from UV exposure. On the other hand, sealants which lack sufficient UV stability, generally show little dirt pick-up, since any dirt deposited on the surface will be washed away as part of the surface erosion caused by the UV degradation. Since our current accelerated weathering machines do not simulate the simultaneous exposure to all weathering factors relevant for inducing surface changes in sealants, we should not expect any correlation of the results obtained in accelerated weathering with those of natural weathering. Furthermore, even natural weathering differs between exposure sites and from year to year, and, therefore, one should not expect reproducibility of the results obtained at different exposure sites at different times. Testing in the laboratory has the benefit of being more reproducible, since it is being carried out with a limited number of variables which can be controlled. The experimenter, thus, is facing the classical dilemma: the laboratory test is more reproducible, but does not consider all relevant weathering factors, the outdoor exposure test comprises all weathering factors active for the specific exposure site and exposure time, but is considerably less reproducible. What, then, is the value of internationally standardising the test method? The benefit is that a test standard defining experimental conditions based on best practices results in better comparability of the relative performance of sealants for a given property, in this case, surface changes with ageing. 5.2 Staining of porous substrates (fluid migration) The proposed test standard is based on the recently published ASTM Test Method for Staining of Porous Substances by Joint Sealants (C1248–93). The test consists of subjecting cured standard joint specimens (tensile adhesion joints of dimensions 12x12x50 mm3), under compression according to the manufacturer’s rated movement capability, separately to heat, UV light with alternating water condensation, and ambient exposure conditions for a total of one month. After this conditioning, the specimens are inspected for any signs of fluid migration on the surface and through the cross-section of the substrate. Since micro-porous


marble has been found to be most sensitive to staining by fluid migration, this substrate is recommended when screening the staining potential of sealants. However, the test method can be used with any other porous substrate (such as natural stones or concrete) and may serve to qualify sealants for specific job site substrates. The main limitation of this method and of all other currently known test methods for assessing the long term staining potential of sealants is that they do not detect delayed staining caused by slowly migrating fluids [20]. High molecular weight plasticizers and unreacted polymers, which take considerable time to migrate from the cured sealant, can cause staining in actual service joints which occurs only one or two years after installation. Efforts to detect to the migration of these high molecular weight fluids in accelerated test methods by extending the duration of the 70°C heat storage to 106 days have been unsuccessful [20]. On the other hand, sealants that showed staining in the proposed accelerated test method were shown to have a high probability of causing staining in actual service joints [1]. The proposed test method can therefore serve as a good screening test, eliminating the majority of sealants that cause problems on projects. Another key aspect of the test method is that it evaluates a specific sealant/ substrate combination. Since, for instance, natural stones show a great variability in micro-porosity between different sources and lots, the quantitative results obtained by the test method, e.g. a certain number of millimetres of staining on a carrerra marble substrate, can only be considered correct for the specific substrate sample evaluated in the test. If a stone from a different source or lot (even colour) is used on the job site than in the laboratory test, the results of the laboratory test may not be applicable. Despite these limitations, the test method can serve for a number of purposes: first, it allows to compare the staining potential of various sealants on the same substrate under identical conditions, second, it allows a sealant manufacturer to qualify a sealant for a specific job site substrate, and finally, as historical data start to accumulate, it identifies and highlights sealants with a high staining potential and, thus, induces the development of less staining sealants. 6 Summary The appearance of any building changes with time as a result of complex processes generally summarised under the term “ageing”. While sealants can contribute to these processes, many aesthetic issues will appear on a building facade over time with or without the presence of a sealant. A work group (WG10) has been charted within the International Standardisation Committee TC59/SC8 (Building Construction, Jointing Products) to assess the need for developing test standards on aesthetic issues caused by sealants. Two work items are being pursued for standardisation: “Staining of


Porous Substrates” and “Determination of Surface Changes Following Exposure to Artificial or Natural Weathering”. The benefits and limitations of the proposed test methods were discussed. 7 References 1.

2. 3. 4. 5.

6. 7. 8.

9.

10.

11.

12.

13.

14.

Wolf, A.T. (1999) Environmental effects on aesthetics of building sealants and adjacent substrates, in Durability of Building Sealants—A State-of-thc-Art Report, (ed. A.T. Wolf), RILEM Publications, Paris (in preparation). Anon. (1992) JSTM J7601T Test Method of Outdoor Exposure for Dirt Collection on Exterior Building Wall Materials and Finishes, Tokyo, Japan. Anon. (1990) Accelerated Dirt Tester, Suga Test Instruments Co., Ltd., Tokyo, Japan. Anon. (1992) JSTM J7602T Test Method for Accelerated Dirt Collection on Exterior Building Wall Materials and Finishes, Tokyo, Japan. O’Neil, V.K. and Altes, M.G. (1994) Evaluation of the fungal resistance of sealant samples under various outdoor exposure conditions, Dow Corning Internal Research Report, Dow Corning Corporation, Midland, Michigan, USA. Anon. (1993), Test Method for Staining of Porous Substances by Joint Sealants (C-1248), American Society for Testing and Materials, Philadelphia. Robinson, R. (1995) Results of round-robin test, communicated at ISO TC59/SC8 meeting, Paris. O’Neil, V.K. and Wolf, A.T. (1996) Effects of weather-proofing sealants on building aesthetics—part 2., in Science and Technology of Building Seals, Sealants, Glazing and Waterproofing: Fifth Volume, ASTM STP 1271, (ed. M.A. Lacasse), American Society for Testing and Materials, Philadelphia, pp. 141–168. Wolf, A.T. (1981) Unpublished results on outdoor weathering of various polysulfide and polyurethane sealant formulations, Perennator GmbH, Wiesbaden, Germany. Wolf, A.T. (1987) Unpublished inspection reports of various buildings stained by putties, butyl caulks, polysulfide and acrylic sealants, Dow Corning S.A., Seneffe, Belgium. Woolman, R. (1994) Staining of marble, in Resealing of Buildings—A Guide to Good Practice, (Ed. Allan Hutchinson), Butterworth Heinemann, London, pp. 117–119. O’Neil, V., Klosowski, J, Altes, M. and Bergman, L. (1994) Effects of weatherproofing sealants on building aesthetics, in Science and Technology of Building Seals, Sealants, Glazing and Waterproofing: Fourth Volume, ASTM STP 1243, (ed. D. Nicastro), American Society for Testing and Materials, Philadelphia, pp. 79–89. Farmer, M.C. and Cechner, R.A. (1996) Laboratory testing of sealants with a marble Substrate, in Science and Technology of Building Seals, Sealants, Glazing and Water-proofing: Fifth Volume, ASTM STP 1271, (ed. M.A. Lacasse), American Society for Testing and Materials, Philadelphia, pp. 184–206. Chin, I.R., Gorrell, T.A. and Scheffler, M.J. (1996) Staining potential of sealants in/ on exterior wall substrates, in Science and Technology of Building Seals, Sealants,


15. 16.

17.

18.

19.

20.

Glazing and Waterproofing: Fifth Volume, ASTM STP 1271, (ed. M.A. Lacasse), American Society for Testing and Materials, Philadelphia, pp. 169–183. Anon. (1974), ISO 2810 Paints and varnishes—Notes for guidance on the conduct of natural weathering test, International Standardisation Organisation, Geneva. Anon. (1994), ISO 2813 Paints and varnishes—Determination of specular gloss of non-metallic paint films at 20°, 60° and 85°, International Standardisation Organisation, Geneva. Anon. (1984), ISO 7724 Paints and varnishes—Colorimetry—Part 1: Principles, Part 2: Colour measurement, Part 3: Calculation of colour difference, International Standardisation Organisation, Geneva. Anon. (1983), ISO 4628 Paints and varnishes—Evaluation of degradation of paint coatings—Designation of intensity, quantity and size of common types of defect, International Standardisation Organisation, Geneva. Fedor, G.R. (1992) Usefulness of accelerated test methods for sealant weathering, in Science and Technology of Building Seals, Sealants, Glazing and Waterproofing: Second Volume, ASTM STP 1200, (ed. J.M. Klosowski), American Society for Testing and Materials, Philadelphia, pp. 48–50. Snyder, R., Badour, R., Carbary, L.D. and Wolf, A.T. (1998) Comparison of various test methods for assessing the long-term fluid migration potential of sealants, in Science and Tcchnology of Building Seals, Sealants, Glazing and Waterproofing: Seventh Volume, ASTM STP 1334, (ed. J.M. Klosowski), American Society for Testing and Materials, Philadelphia (in print).

Index

This index is based on keywords assigned to the individual chapter. The numbers are the page numbers of the first page of the relevant chapter. A Accelerated ageing, 47 Accelerated weathering, 171 Adhesion, 23, 47, 161 Ageing, 1, 31, 151 Aluminium, 99, 117

Dynamic mechanical analysis, 81 E Elongation, 81 Engineering, 63 F Face-sealed joints, 133 Fatigue, 161 Fatigue analysis, 23 Field inspection, 133 Fluid migration, 181 Fracture, 161 Functional requirements, 133

B Bead dimension. 63 Building, 99 C Carbon flamc, 171 Chalking, 181 Chemical analysis, 151 Cladding, 99 Cladding systems, 133 Composition, 1 Concrete, 99, 117 Crazing, 181 Crosslink, 1 Curing, 117 Cyclic movement, 31

G Glass transition temperature, 81 H Hardening 1 I Immersion, 1 Interphase, 161

D Degradation, 151 Durability, 47, 161

J Joint, 99 211

212 INDEX

M Mechanical properties, 81, 117 Modulus of elasticity, 81 Movement, 99 Movement during cure, 23, 99, 117 Movements rig, 117 O OEM window fabrication, 23 Outdoor exposure, 171 P Polysulfide, 151 Polyurethane, 151 Q QUV, 171 R Re-seal jointing systems, 133 Rheometry, 81 S Sealant, 23, 31, 63, 99, 151, 161, 181 Sealant joint performance, 117 Sealants,1 , 47, 81 Service life, 133 Silicone, 151 Silicone sealant, 171 Silicones, 1 Specifications, 133 Staining, 181 Strength, 81 Structural glazing, 47 Structural silicone, 63 Surface changes, 181 T Testing, 47 Thermal analysis, 151 Time-compression factor, 31 V Viscoelasticity, 161 W Weathering, 1, 31

X Xenon lamp, 171

Durability of Building Sealants (Rilem Proceedings, 36)

Durability of Geotextiles (RILEM Proceedings)

Durability of Building and Construction Sealants and Adhesives: 2nd Volume

Durability Design of Concrete Structures (Rilem Report)

Vegetable Plants and their Fibres as Building Materials: Proceedings of the Second International RILEM Symposium (RILEM Proceedings 7)

Admixtures for Concrete - Improvement of Properties: Proceedings of the International RILEM Symposium (RILEM Proceedings 5)

Test Quality for Construction, Materials and Structures: Proceedings of the International RILEM ILAC Symposium (RILEM Proceedings)

Performance Criteria for Concrete Durability (Rilem Report, No 12)

Special Concretes - Workability and Mixing (Rilem Proceedings)

Durability of Building Materials and Components 8: Service life and durability of materials and components

Durability of Strain-Hardening Fibre-Reinforced Cement-Based Composites (SHCC) (RILEM State of the Art Reports)

Durability of Materials and Structures in Building and Civil

Durability of Strain-Hardening Fibre-Reinforced Cement-Based Composites (SHCC) (RILEM State of the Art Reports)

Handbook of Adhesives and Sealants

Durability of Building Materials and Components: Proceedings of the 5th International Conference Held in Brighton, UK, 7-9 November 1990

Reinforced Plastics Durability

Reinforced Plastics Durability

Adhesives and Sealants

Performance and Durability of Bituminous Materials: Proceedings of Symposium, University of Leeds, March 1994

Adhesives & Sealants Industry November 2011

36 Hours

36 Exposures

36 Stunden

36 Children

Week 36

36 Hours

Reality 36

36 Exposures

36 Stunden

99 (36)

Durability of Building Sealants (Rilem Proceedings, 36)

Durability of Geotextiles (RILEM Proceedings)

Durability of Building and Construction Sealants and Adhesives: 2nd Volume

Durability Design of Concrete Structures (Rilem Report)

Vegetable Plants and their Fibres as Building Materials: Proceedings of the Second International RILEM Symposium (RILEM Proceedings 7)

Admixtures for Concrete - Improvement of Properties: Proceedings of the International RILEM Symposium (RILEM Proceedings 5)

Test Quality for Construction, Materials and Structures: Proceedings of the International RILEM ILAC Symposium (RILEM Proceedings)

Performance Criteria for Concrete Durability (Rilem Report, No 12)

Special Concretes - Workability and Mixing (Rilem Proceedings)

Durability of Building Materials and Components 8: Service life and durability of materials and components

Durability of Strain-Hardening Fibre-Reinforced Cement-Based Composites (SHCC) (RILEM State of the Art Reports)

Durability of Materials and Structures in Building and Civil

Durability of Strain-Hardening Fibre-Reinforced Cement-Based Composites (SHCC) (RILEM State of the Art Reports)

Handbook of Adhesives and Sealants

Durability of Building Materials and Components: Proceedings of the 5th International Conference Held in Brighton, UK, 7-9 November 1990

Reinforced Plastics Durability

Reinforced Plastics Durability

Adhesives and Sealants

Performance and Durability of Bituminous Materials: Proceedings of Symposium, University of Leeds, March 1994

Adhesives & Sealants Industry November 2011

36 Hours

36 Exposures

36 Stunden

36 Children

Week 36

36 Hours

Reality 36

36 Exposures

36 Stunden

99 (36)

Recommend Documents