Microengineering of Metals and Ceramics: Part I: Design, Tooling, and Injection Molding. Part II: Special Replication Techniques, Automation, and Properties (Advanced Micro and Nanosystems)

Advanced Micro & Nanosystems Volume 3 Microengineering of Metals and Ceramics

Part I Design, Tooling and Injection Molding

Advanced Micro and Nanosystems Vol. 3. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31208-0

Related Titles Other AMN Volumes Baltes, H., Brand, O., Fedder, G. K., Hierold, C. Korvink, J. G., Tabata, O. (eds.)

Champion, Y., Fecht, H.-J. (eds.)

Enabling Technologies for MEMS and Nanodevices

Fabrication, Control and Properties

Advanced Micro and Nanosystems

Nano-Architectured and Nanostructured Materials 2004 Hardcover

2004

ISBN 3-527-31008-8

Hardcover ISBN 3-527-30746-X

Wagner, L. (ed.) Baltes, H., Brand, O., Fedder, G. K., Hierold, C., Korvink, J. G., Tabata, O. (eds.)

CMOS-MEMS

Shot Peening 2003 Hardcover ISBN 3-527-30537-8

2005 Hardcover ISBN 3-527-31080-0

Löhe, D., Haußelt, J. (eds.)

Microengineering of Metals and Ceramics Part II: Special Replication Techniques, Automation and Properties 2005 Hardcover ISBN 3-527-31493-8

Scheffler, M., Colombo, P. (eds.)

Cellular Ceramics Structure, Manufacturing, Properties and Applications 2005 Hardcover ISBN 3-527-31320-6

Schulze, V.

Modern Mechanical Surface Treatment States, Stability, Effects

Kockmann, N. (ed.)

Micro Process Engineering Fundamentals, Devices, Fabrication, and Applications 2006 Hardcover ISBN 3-527-31246-3

2005 Hardcover ISBN 3-527-31371-0

Advanced Micro & Nanosystems Volume 3 Microengineering of Metals and Ceramics Part I Design, Tooling and Injection Molding Volume Editors Detlef Löhe and Jürgen Haußelt

Volume Editors Prof. Dr.-Ing. Detlef Löhe Institute. of Materials Science and Engineering Universität Karlsruhe Kaiserstr. 12 76131 Karlsruhe Germany Prof. Dr. Jürgen Haußelt Institute for Materials Research III Forschungszentrum Karlsruhe Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen Germany

n All books published by Wiley-VCH are carefully produced. Nevertheless, authors, editors, and publisher do not warrant the information contained in these books, including this book, to be free of errors. Readers are advised to keep in mind that statements, data, illustrations, procedural details or other items may inadvertently be inaccurate.

Cover: Explosion drawing by Tobias Deigendesch and Jochen Marz

Series Editors Prof. Henry Baltes Physical Electronics Laboratory ETH Zürich 8093 Zürich Switzerland Oliver Brand School of Electrical and Computer Engineering Georgia Institute of Technology 777 Atlantic Drive Atlanta, GA 30332-0250 USA Prof. Dr. Gary K. Fedder ECE Department & Robotics Institute Carnegie Mellon University Pittsburgh, PA 15213-3890 USA Prof. Dr. Christofer Hierold Chair of Micro- and Nanosystems ETH Zürich ETH-Zentrum, CLA H9 Tannenstr. 3 8092 Zürich Switzerland Prof. Dr. Jan G. Korvink Institute for Microsystem Technology (IMTEK) Albert-Ludwigs-Universität Freiburg Georges-Köhler-Allee 103 79110 Freiburg Germany Prof. Dr. Osamu Tabata Dept. of Mechanical Engineering Kyoto University Yoshida Honmachi Sakyo-ku 606-8501 Kyoto Japan

Library of Congress Card No.: applied for British Library Cataloguing-in-Publication Data: A catalogue record for this book is available from the British Library. Bibliographic information published by Die Deutsche Bibliothek Die Deutsche Bibliothek lists this publication in the Deutsche Nationalbibliografie; detailed bibliographic data is available in the Internet at © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim, Germany All rights reserved (including those of translation in other languages). No part of this book may be reproduced in any form – by photoprinting, microfilm, or any other means – nor transmitted or translated into machine language without written permission from the publishers. Registered names, trademarks, etc. used in this book, even when not specifically marked as such, are not to be considered unprotected by law. Typesetting K+V Fotosatz GmbH, Beerfelden Printing betz-Druck GmbH, Darmstadt Bookbinding J. Schäffer GmbH, Grünstadt Printed in the Federal Republic of Germany Printed on acid-free paper ISBN-13: ISBN-10:

978-3-527-31208-5 3-527-31208-0

V

Preface Machines and their design and production have fascinated mankind from the very beginnings of culture. The last decades have shown mechanical contraptions decreasing in size to almost invisible dimensions. The implementation of micromechanics has become not only a technological challenge, but also a necessity for a successful future development of whole industrial branches. Adequate design and replication techniques of micromechanical components as well as a deep knowledge of their properties are indispensable for further progress in this field. At the same time the variety of materials used in micro system technology has increased significantly. Today not only silicon and polymers, but also metals and ceramics are of increasing interest for a large number of applications. In contrast to silicon and polymers, however, which can be structured by technologies well known from microelectronics, metals and ceramics require new forming and structuring techniques for dimensions in the sub millimeter range. In addition, mechanical properties of metallic and ceramic microparts are of special interest because they differ significantly from those measured in macroscopic dimensions. It is because of these considerations that the German Research Council (Deutsche Forschungsgemeinschaft, DFG) has decided to fund a collaborative research center (SFB 499) with approximately 2 million euros p.a. which tackles the problems arising when trying to design, produce and characterize advanced microstructures made of metals and ceramics. The two-volume book in hand presents the results of five years of research on micro engineering utilizing metallic and ceramic materials. It comprises the whole process chain from design and modeling of microcomponents along production preparation and two central replication techniques (micro powder injection molding and micro casting) to characterization and quality insurance, the scope encompassing both theoretical and experimental topics. The book is structured roughly according to the project groups which form SFB 499: The first volume contains sections on design, tooling and replication techniques based on injection molding. The first section focuses on micro-component design including design environment, design flow, modeling and validation as well as on the modeling of micro powder injection molding (cf. chapters 1 to 3). Advanced Micro and Nanosystems Vol. 3. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31208-0

VI

Preface

The second section on tooling describes preparatory steps for the production process. The production of mold inserts by micro milling, laser ablation, micro electro discharge machining and techniques based on lithography and electroforming is investigated, surface treatment methods using shot peening and ultrasonic energy are presented, and optimized mold materials for micro casting are identified (cf. chapters 4 to 9). The third section on micro injection molding (chapters 10 to 12) is concerned with the production itself, focusing on the actual molding processes. Following an introductory chapter on general aspects of micro injection molding, micro injection molding of metals and ceramics including the challenging process steps of debinding and sintering (cf. chapters 11 and 12) are described. The second volume comprises three sections on replication techniques other than injection molding, on automation and on properties of the components produced. The section on special replication techniques focuses on microcasting (chapter 13) and electroforming (chapter 14) of metals and on selected techniques for the manufacturing of ceramic microcomponents (chapter 15). The second section of volume 2 deals with automation and quality insurance and includes chapters on the automation of lPIM, on assembly and on quality insurance and dimensional measuring techniques (chapters 16 to 18). The last section addresses the properties of metals and ceramics and of the components produced. Microstructure and mechanical properties including micromechanical testing under quasi-static and cyclic loading as well as tribology are investigated and numerical wear simulation is performed (cf. chapters 19 to 22). Working groups concerned with aspects touching all five project sections (e.g. on relationships between manufacturing processes) act as links between the projects. Cooperation within SFB 499 is further enforced by concerted work on a demonstration device consisting of a micro-turbine and a sun-and-planet gearing. For a comprehensive treatment the chapters which directly deal with the research projects of SFB 499 are supplemented by several important research topics concerning micromechanical components (e.g. laser structuring, lithographic processes, electroforming, assembly) which are not part of the collaborative research center. In these cases guest authors have been asked to contribute. They come from Bremen University (Institut für angewandte Strahltechnik, BIAS), from Braunschweig University (Institut für Werkzeugmaschinen und Fertigungstechnik, IWF) and several scientific institutes of Forschungszentrum Karlsruhe. By covering most aspects of the design, production, and properties of micromechanical components outside the silicon world, the authors hope to present a useful guide to students and readers looking for a comprehensive overview as a starting point of in-depth research in this field. However, the detailed presentation of latest SFB 499 research results as well as contributions from literature should also be a source of new insights and inspiration for micro-engineering experts from research institutions and industry. May 2005, Karlsruhe

Detlef Löhe and Jürgen Haußelt, Volume Editors

VII

Foreword We are proud to present the third and fourth volumes of Advanced Micro & Nanosystems (AMN), entitled Microengineering of Metals and Ceramics. Although microtechnology is often associated with semiconductor cleanroom processes, this is by no means the only means of production available. The processes we associate with traditional mechanical engineering mass production have also been the focus of microtechnologists, with tremendous successes already in place and a huge potential for further progress. Of course, every new technology pairs the development of suitable materials with that of production technology, and in the sub-millimeter range the challenges become immense. Not only must raw materials be produced in particulate form fine enough to reproduce the molds they are formed into, and molds need to be prepared at the correct dimensions and surface quality, but new ideas are needed to make use of machine parts produced in this manner, and new methods to assemble parts into complete systems. In these two volumes you will find a comprehensive treatment of a variety of challenges that arise in the process of producing microparts from metals and ceramics, from materials, testing, production, computer aided engineering all the way to assembly. We hope that these volumes will inspire the transfer of these fascinating techniques not only to other research groups, but also to industry and so broaden the range of items that can be successfully miniaturized. Covering recent advances from the world of micro and nanosystems, future AMN issues will either focus on a particular subject, such as CMOS-MEMS and the present twin topical volumes Microengineering of Metals and Ceramics, or be a carefully chosen set of cutting-edge overview and review articles like the first AMN volume on Enabling Techniques for MEMS and Nanodevices. Looking ahead, we hope to welcome you back, dear reader, to the upcoming fifth member of the AMN series, in which we take a close look at the fascinating field of Micro Process Engineering. The articles will range from the fundamentals and engineering over device conception and simulation to fabrication strategies and techniques, and finally cover application and operational issues. To cover such a wide spectrum, we are very glad to have the support of Dr. NorAdvanced Micro and Nanosystems Vol. 3. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31208-0

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Foreword

bert Kockmann from the University of Freiburg, Germany, who will edit this volume. Henry Baltes, Oliver Brand, Gary K. Fedder, Christofer Hierold, Jan G. Korvink, and Osamu Tabata Series Editors May 2005 Zurich, Atlanta, Pittsburgh, Freiburg and Kyoto

IX

Contents Preface

V

Foreword VII List of Contributors

XI

I

Design

1

Design Environment and Design Flow

3

A. Albers, J. Marz 2

Modeling and Validation in Design

29

A. Albers, D. Metz 3

Modeling Micro PIM 51

D. Kauzlaric, A. Greiner, J. G. Korvink, M. Schulz, R. Heldele II

Tooling

4

Strategies for the Manufacture of Mold Inserts 87

J. Fleischer, C. Buchholz, J. Kotschenreuther 5

Micro End Milling of Hardened Steel 107

J. Schmidt, J. Kotschenreuther 6

3D Microstructuring of Mold Inserts by Laser-based Removal

A. Stephen, F. Vollertsen 7

Micro-EDM for Mold Inserts

161

J. Schmidt, M. Knoll, T. Masuzawa Advanced Micro and Nanosystems Vol. 3. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31208-0

131

X

Contents

8

Lithographic Fabrication of Mold Inserts 187

M. Guttmann, J. Schulz, V. Saile 9

Material States and Surface Conditioning for Mold Inserts

221

Ch. Horsch, V. Schulze, D. Löhe

III

Replication Techniques – Microinjection Molding

10 Microinjection Molding – Principles and Challenges

253

R. Ruprecht, G. Finnah, V. Piotter 11 Micro Metal Injection Molding

289

V. Piotter, L. Merz, G. Örlygsson, S. Rath, R. Ruprecht, B. Zeep 12 Micro Ceramic Injection Molding

325

W. Bauer, J. Hausselt, L. Merz, M. Müller, G. Örlygsson, S. Rath Subject Index

XVII, end of the book

XI

List of Contributors Prof. A. Albers Institut für Produktentwicklung Universität Karlsruhe Kaiserstrasse 12 76128 Karlsruhe Germany

T. Beck Institut für Werkstoffkunde I Universität Karlsruhe Kaiserstrasse 12 76131 Karlsruhe Germany

M. Auhorn Institut für Werkstoffkunde I Universität Karlsruhe Kaiserstrasse 12 76131 Karlsruhe Germany

I. Behrens Institut für Werkzeugmaschinen und Betriebstechnik Universität Karlsruhe Kaiserstrasse 12 76131 Karlsruhe Germany

K. Bade Forschungszentrum Karlsruhe Institut für Mikrostrukturtechnik Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen Germany W. Bauer Forschungszentrum Karlsruhe Institut für Materialforschung III Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen Germany G. Baumeister Forschungszentrum Karlsruhe Institut für Materialforschung III Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen Germany

H. von Both Institut für Mikrosystemtechnik Albert-Ludwigs-Universität Freiburg Georges-Köhler-Allee 103 79110 Freiburg Germany C. Buchholz Institut für Werkzeugmaschinen und Betriebstechnik Universität Karlsruhe Kaiserstrasse 12 76131 Karlsruhe Germany M. Dauscher Forschungszentrum Karlsruhe Institut für Materialforschung III Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen Germany

Advanced Micro and Nanosystems Vol. 3. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31208-0

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List of Contributors

G. Finnah Forschungszentrum Karlsruhe Institut für Materialforschung III Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen Germany

R. Heldele Forschungszentrum Karlsruhe Institut für Materialforschung III Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen Germany

Prof. J. Fleischer Institut für Werkzeugmaschinen und Betriebstechnik Universität Karlsruhe Kaiserstrasse 12 76131 Karlsruhe Germany

J. Herz Institut für Werkstoffkunde II Universität Karlsruhe c/o Forschungszentrum Karlsruhe Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen Germany

Prof. K.-H. zum Gahr Institut für Werkstoffkunde II Universität Karlsruhe c/o Forschungszentrum Karlsruhe Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen Germany

Prof. J. Hesselbach Institut für Werkzeugmaschinenund Fertigungstechnik Technische Universität Braunschweig Langer Kamp 19 b 38106 Braunschweig Germany

A. Greiner Institut für Mikrosystemtechnik Albert-Ludwigs-Universität Freiburg Georges-Köhler-Allee 10 379110 Freiburg Germany M. Guttmann Forschungszentrum Karlsruhe Institut für Mikrostrukturtechnik Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen Germany Prof. J. Haußelt Forschungszentrum Karlsruhe Institut für Materialforschung III Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen Germany V. Hegadekatte Institut für Zuverlässigkeit von Bauteilen und Systemen Universität Karlsruhe Kaiserstrasse 12 76131 Karlsruhe Germany

K. Heuer Institut für Werkzeugmaschinen und Fertigungstechnik Technische Universität Braunschweig Langer Kamp 19b 38106 Braunschweig Germany C. Horsch Institut für Werkstoffkunde I Universität Karlsruhe Kaiserstrasse 12 76131 Karlsruhe Germany N. Huber Institut für Materialforschung II Forschungszentrum Karlsruhe Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen Germany B. Kasanická Institut für Werkstoffkunde I Universität Karlsruhe Kaiserstrasse 12 76131 Karlsruhe Germany

List of Contributors

D. Kauslaric Institut für Mikrosystemtechnik Albert-Ludwigs-Universität Freiburg Georges-Köhler-Allee 103 79110 Freiburg Germany

J. Marz Institut für Produktentwicklung Universität Karlsruhe Kaiserstrasse 12 76128 Karlsruhe Germany

R. Knitter Forschungszentrum Karlsruhe Institut für Materialforschung III Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen Germany

T. Masuzawa Institute of Industrial Science University of Tokyo 4-6-1 Komaba, Meguro-ku Tokyo Japan

M. Knoll Institut für Werkzeugmaschinen und Betriebstechnik Universität Karlsruhe Kaiserstrasse 12 76131 Karlsruhe Germany

L. Merz Forschungszentrum Karlsruhe Institut für Materialforschung III Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen Germany

Prof. J. G. Korvink Institut für Mikrosystemtechnik Albert-Ludwigs-Universität Freiburg Georges-Köhler-Allee 103 79110 Freiburg Germany J. Kotschenreuther Institut für Werkzeugmaschinen und Betriebstechnik Universität Karlsruhe Kaiserstrasse 12 76131 Karlsruhe Germany O. Kraft Institut für Zuverlässigkeit von Bauteilen und Systemen Universität Karlsruhe Kaiserstrasse 12 76131 Karlsruhe Germany Prof. D. Löhe Institut für Werkstoffkunde I Universität Karlsruhe Kaiserstrasse 12 76131 Karlsruhe Germany

D. Metz Institut für Produktentwicklung Universität Karlsruhe Kaiserstrasse 12 76128 Karlsruhe Germany M. Müller Forschungszentrum Karlsruhe Institut für Materialforschung III Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen Germany G. Örlygsson Forschungszentrum Karlsruhe Institut für Materialforschung III Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen Germany V. Piotter Forschungszentrum Karlsruhe Institut für Materialforschung III Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen Germany

XIII

XIV

List of Contributors

A. Raatz Institut für Werkzeugmaschinen und Fertigungstechnik Technische Universität Braunschweig Langer Kamp 19 b 38106 Braunschweig Germany

J. Schneider Institut für Werkstoffkunde II Universität Karlsruhe c/o Forschungszentrum Karlsruhe Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen Germany

S. Rath Forschungszentrum Karlsruhe Institut für Materialforschung III Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen Germany

J. Schulz Forschungszentrum Karlsruhe Institut für Mikrostrukturtechnik Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen Germany

H.-J. Ritzhaupt-Kleissl Forschungszentrum Karlsruhe Institut für Materialforschung III Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen Germany

M. Schulz Forschungszentrum Karlsruhe Institut für Materialforschung III Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen Germany

R. Ruprecht Forschungszentrum Karlsruhe Institut für Materialforschung III Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen Germany

V. Schulze Institut für Werkstoffkunde I Universität Karlsruhe Kaiserstrasse 12 76131 Karlsruhe Germany

Prof. V. Saile Forschungszentrum Karlsruhe Institut für Mikrostrukturtechnik Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen

S. Soetebier Institut für Werkzeugmaschinen und Fertigungstechnik Technische Universität Braunschweig Langer Kamp 19 b 38106 Braunschweig Germany

G. Schanz Forschungszentrum Karlsruhe Institut für Materialforschung III Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen Germany Prof. J. Schmidt Institut für Werkzeugmaschinen und Betriebstechnik Universität Karlsruhe Kaiserstrasse 12 76131 Karlsruhe Germany

A. Stephen Bremer Institut für angewandte Strahltechnik (BIAS) Klagenfurter Str. 2 28359 Bremen Germany Prof. F. Vollertsen Bremer Institut für angewandte Strahltechnik (BIAS) Klagenfurter Strasse 2 28359 Bremen Germany

List of Contributors

H. Weule Institut für Werkzeugmaschinen und Betriebstechnik Universität Karlsruhe Kaiserstrasse 12 76131 Karlsruhe Germany J. Wrege Institut für Werkzeugmaschinen und Fertigungstechnik Technische Universität Braunschweig Langer Kamp 19 b 38106 Braunschweig Germany

B. Zeep Forschungszentrum Karlsruhe Institut für Materialforschung III Hermann-von-Helmholtz-Platz 1 76344 Eggenstein-Leopoldshafen Germany

XV

I Design

Advanced Micro and Nanosystems Vol. 3. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31208-0

3

1 Design Environment and Design Flow A. Albers, J. Marz, Institute of Product Development (IPEK), University of Karlsruhe (TH), Germany Abstract The design flow for primary-shaped microcomponents and microsystems is presented. As a characteristic of microspecific design, the approach is predominantly driven by technology. To integrate the relevant technological demands and restrictions into the design synthesis for a realizable embodiment design in accordance with the specified function, design rules are defined. These represent mandatory instructions for the designer. To support the designer effectively the design rules are provided within a computer-aided design environment. In addition to an information portal, an embodiment design unit is built up on the basis of the 3D CAD system Unigraphics, which includes an application for knowledge-based engineering (KBE). The rule-based design methodology was used for the development and design of a microplanetary gear.

Keywords design environment; design flow; target system definition; operation system; object system; design rule; knowledge-based engineering; methodological aid

1.1

Introduction

1.1.1

State-of-the-Art of Design Flows and Design Environments within Microtechnology 4 Mechanical Microproduction 5

1.1.2 1.2 1.2.1 1.2.1.1 1.2.1.2 1.2.1.3 1.2.1.4 1.2.1.5 1.2.1.6

Design Flow

4

6

Specific Issues Within the Design of Microsystems 6 Dominance of Technologies 6 Surface-to-Volume Ratio 6 Dynamics 7 Standardization 7 Validation 7 Enhanced Material Spectrum 7

Advanced Micro and Nanosystems Vol. 3. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31208-0

4

1 Design Environment and Design Flow

1.2.1.7 1.2.2 1.3 1.3.1 1.3.1.1 1.3.1.2 1.3.2 1.3.2.1 1.3.2.2 1.4 1.4.1 1.4.2 1.4.2.1 1.4.2.2 1.5 1.6

Emphasis on Actuators 7 Microspecific Design Flow 7 Design Rules 9

Basics 9 Definition 9 Derivation of Design Rules 10 Design Rules Derived from Restrictions of Production Technology 11 Design Rules for Mold Insert Manufacturing 13 Design Rules for Replication Techniques 14 Design Environment

18

Information Unit 20 Embodiment Design Unit 20 Preparing Elementary Rules for Computer-aided Design Rule Check 21 Design Rule Check 24 Conclusion 27 References 27

1.1 Introduction

Microtechnology involves technologies for manufacturing and assembling predominantly micromechanical, microelectrical, microfluidic and microoptical components and systems with characteristic structures with the dimension of microns. In doing so, microproduction technologies take on a key role, since their process-specific parameters and boundary conditions determine the smallness and attainable quality features of the components. Owing to the ongoing progress in microtechnology and the increasing penetration of the market with medium-sized and large-batch products, development steps preliminary and subsequent to production are becoming more and more relevant for an effective design in compliance with the requirements. Therefore, the designer needs to be supported by a technological basic knowledge and know-how, regardless of individual persons. 1.1.1 State-of-the-Art of Design Flows and Design Environments within Microtechnology

Microtechnologies include silicon microsystem technology, the LIGA process and mechanical microproduction technology. Silicon microsystem technology is the most widespread microtechnology throughout the world. It is based on the process technology of integrated circuits (ICs) and benefits from a comprehensive know-how from microelectronics. Unlike in microelectronics, microtechnological products integrate active

1.1 Introduction

and passive functional elements, which rely on at least two elementarily different physical, chemical or biochemical effects and working principles. In addition to sensors and information processing, particularly actuator functions are performed. The predominantly 2.5-dimensional and sometimes three-dimensional structures use silicon as substrate with its excellent mechanical properties. Along with others, all these characteristics of silicon micromechanical systems have required a specific design methodology ever since a critical level of development from research into industry was reached. Different design process models are known [1–3], which among other things integrate analytical and numerical simulation tools. Silicon-based micromechanical products are developed in an iterative sequence of synthesis and analysis steps. A specific difficulty lies in the deviation between the designed target structure and the actual structure after the optical lithography and etching process. Therefore, compensation structures are introduced into the design and simulation environment, adjusting the determined structure by dimensional add-on and auxiliary structures [4–6]. Design rules are introduced as a methodological aid to represent this technological information. Design rules have been used in microelectronics since the early 1980s to enable very large-scale integrated (VLSI) circuits to be synthesized automatically to the extent of nearly 100% [7]. Silicon microsystem technology has now reached a high degree of development status. A lot of research programs have led to design flow descriptions and collections of design rules. Like silicon microsystem technology, the LIGA process utilizes mask-based process steps. The LIGA process approaches an obviously broader range of materials and is characterized by extremely high aspect ratios with at the same time the smallest lateral structure dimensions [8]. LIGA permits the manufacture of mold inserts which can be used in replication techniques for large-batch parts (Chapter 8). In addition to thermoplastics, also metallic and ceramic materials are processed. To support the design and process, engineering design rules are utilized which give – depending on the process sequence – instructions for a design for manufacturing and for separating, manipulating and assembling components [9]. Within different research programs, design environments for computer-aided design of LIGA microstructures embedding design rules were developed [10, 11]. The computer-aided design of LIGA microstructures still shows a high demand. A standardized model for methodological design flow in the LIGA process is lacking to date [12]. 1.1.2 Mechanical Microproduction

To come up with a more cost-effective, medium-sized and large-batch suitable process for manufacturing microsystems, the potential of miniaturizing mechanical production technologies has been increasingly investigated in recent years. Predominantly staged production process sequences for manufacturing mold inserts by wear-resistant materials followed by a replication step show out-

5

6

1 Design Environment and Design Flow

standing future prospects. Technologies such as micromilling and laser machining are suitable for manufacturing complex three-dimensional free-form surfaces (Chapters 5–7). By replication techniques such as micropowder injection molding, high-strength microcomponents and microsystems from metallic and ceramic materials can be produced in large quantities (Chapters 11 and 12). When designing primary-shaped microparts with respect to function and manufacturing, it is necessary to incorporate boundary conditions and restrictions from process steps downstream to the product development into the design activities as early as possible [20]. Thus, a design flow is introduced that uses design rules to support the designer effectively with respect to functional, geometric and capacitance demands. The process model and the method are embedded in a knowledge-based design environment.

1.2 Design Flow 1.2.1 Specific Issues Within the Design of Microsystems

In contrast to the procedures and methods commonly applied in mechanical engineering and precision engineering, product development of microtechnological systems requires attention to the following issues.

1.2.1.1 Dominance of Technologies Going beyond the basic rules and guidelines of embodiment design microtechnology has a strong focus on parallelization of product and process development. Resulting from the rapid advances in existing production processes and the appearance of new technologies, the question of ‘how to manufacture’ becomes a conceptual part of product development. Microproduction technologies, materials and specific effects define the possible shape and function of new products.

1.2.1.2 Surface-to-Volume Ratio Owing to the super-proportional rise in the surface-to-volume ratio in the range of the characteristic and functional dimensions of microcomponents, the global dimensions have a different ratio to local deviances. Higher level surface tolerances in macroengineering have the same significance as notch form deviation in microengineering. There is no longer a difference in magnitude between material microstructures and work-piece dimensions. The numbers of crystals and surface layers are relevant for the calculation of elastic properties.

1.2 Design Flow

1.2.1.3 Dynamics As a consequence of their small volumes, microsystems have lower inertia. They can be operated in higher ranges of frequency and show high dynamics.

1.2.1.4 Standardization Standards with regard to generic or product-specific dimensions do not exist for the design of microcomponents and systems.

1.2.1.5 Validation Mostly, either no equipment for the measurement and testing of microcomponents is available at all or insurmountable physical obstacles occur (size of components, essential accuracy of the measuring equipment). Design can, therefore, only set requirements on what can be verified by means of measurement and with the use of testing equipment. Compared with silicon microsystem technology, the LIGA process and the mechanical microproduction technologies show the following specific differences.

1.2.1.6 Enhanced Material Spectrum Microsystem technologies with replication subprocesses possess an enhanced material spectrum. Totally new applications arise from it, making it necessary to characterize the materials with respect to their microstructures and properties. This is an important input for product development.

1.2.1.7 Emphasis on Actuators Since the LIGA process and mechanical microproduction technologies do not rely on silicon as base material, there is enormous potential to develop actuators using a multitude of effects. Integrated in a superior system or as an integrated self-sufficient microsystem, actuators offer particularly energy and material interfaces to the macroscopic world. A microspecific design methodology has to be directed on methods and processes to calculate and design the relevant interface machine elements. 1.2.2 Microspecific Design Flow

Each design process starts with a definition of the target system. The target system definition is developed with the involvement of the customer and determines requirements and boundary conditions for the product that is to be developed (Fig. 1-1). The target system definition helps to concretize the task and to clarify vague and unexpressed demands on the object system – the subsequent microproduct – prior to the beginning of the design. Along with the customer, a requirements list is

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generated, which describes the target system by quantitative and assessed criteria. To ensure that a fundamental criterion is not forgotten, checklists with main headings exist for drawing up a requirements list [13, 14]. The requirements list represents a dynamic document, which has to be examined continually with respect to up-to-dateness and inconsistencies during the design process. Moreover, the risk exists of specifying the task in an unchallenged or in an overextended way. An unchallenged specification might lead to a product ahead of schedule but without matching the real performance characteristics. On the other hand, an overextended specification might limit the solution space in such a manner that no solution could be developed [3]. For the target system definition of microelectronic circuits, hardware description languages are standardized. The microsystem technology of primary shaping concentrates on energy- and material-converting microsystems with integrated information flow and with single functions from different physical, chemical and biochemical domains, so no formal methods and target system definition languages are available. When conventionally developing products and systems of mechanical engineering and precision engineering [13–16], a conceptional phase would follow, in which basic partial solutions for functionally organized subsystems would be developed and systematically combined to the optimum basic solution with consideration of evaluation techniques. When developing microsystems, the approach is ‘technology driven’. At the same time, the technology term describes all of those scientific disciplines as a whole that contribute to the product development process. This especially applies in production engineering and material sciences. Among material sciences, also research on new or specifically formed

Fig. 1-1 Microspecific design flow

1.3 Design Rules

physical, chemical and biochemical effects has to be itemized. Effects are comprehended as both those which are intentionally used to transfer the target system into the object system by effects and active principles in order to fulfil a function (e.g. shape memory effect) and those which inevitably result from phenomena such as friction and wear. Because of being driven by technology, parallelization of stages of conceptual and embodiment design occurs, which exceeds different levels of abstraction. While making conceptual decisions on system level related to function in a topdown approach, simultaneously structural details conditional on technology are being designed in a bottom-up way. In between, single components are preliminarily drafted (basic design). These structural details can be entirely finalized and annotated with all tolerance data and information relevant for production preparation. Already during the subsequent design stage, a complete component can be constituted in its final shape (detail design). The system comes to the stage of basic design. Eventually the system itself is finalized and refined into a detailed design documentation for transfer to production preparation. In doing so, the approach constantly changes between the view on the complete system and the smallest structural element (’meet-in-the-middle‘) [17], wherein the design space is restricted for the designer through boundary conditions and restrictions of the production processes. However, features that cannot be described as easily as geometric quantities also have an influence. These are characteristics of the materials themselves such as microstructure or mechanical properties and physical, chemical or biochemical effects made accessible by them. The latter can develop into disturbing effects when the dimensions become smaller, they can become less important or even emerge and therefore open up completely new applications. All of these ‘technological’ aspects therefore have to be integrated into the microtechnological design of structures, components and systems [18]. Therefore, it is necessary to make the multi-technological knowledge from the above-mentioned technologies directly available to the designer in the design process. This is achieved via the methodological aid of design rules.

1.3 Design Rules 1.3.1 Basics 1.3.1.1 Definition Design rules are instructions derived from technological restrictions which have to be followed mandatorily for a realizable design. Technologies embrace all processes and methods of production preparation, production and material science including effects which are adjacent or subsequent to the design process. Restrictions describe all boundary conditions, requirements and constraints that influence the design embodiment of the prod-

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uct with respect to the entire product life cycle. A realizable design is a design that is completely specified in detail (CAD–CAM suitable 3D CAD model, drawings) ready for production. Owing to their mandatory character, design rules are an explicit part of the conceptional and embodiment design activities. Disregarding the knowledge about restrictions leads to a design that only inadequately fulfils the function or is even not able to be manufactured, assembled, dimensionally characterized and so on. Design rules are mandatory instructions to be followed by the designer!

1.3.1.2 Derivation of Design Rules Design rules begin by detecting potential influences that a technology could have on the microtechnological design. The features and parameters of this technology are quantifiably taken over in an extraction step. Then these properties will be projected to existing and possible components and system structures and marginal analyses of still realizable manufacturing sizes will be made, i.e. interpreted relevant to design [19, 20]. Only a methodical trained product development engineer can carry out this interpretation and raise technical facts via suitable query techniques (interviewing techniques, e-mail query via special forms, so called ‘technology specification sheets’; see, for example, Fig. 1-5). Now these ‘raw data’ need to be transformed via creativity methods or with experimental knowledge into a methodical knowledge that can be used by the designer. Knowledge from individual disciplines – from the designer’s view this is data and information – is raised to a higher level of knowledge and made available mono-disciplinarily, i.e. from the designer’s point of view (Fig. 1-2). According to a specially developed classification scheme, the interpretations are formulated as generally applicable rules. The classification scheme that is introduced here models itself on the technologies for mold insert production and replication that are a part in the production process. The nomenclature is as as shown in Fig. 1-3. These different process steps are formulated separately for the further application of the design rules in a knowledge-based design environment. It has to be clearly determined to which type of part the geometric sizes refer. This is carried out via a so-called rule class. It indicates for which type of part the rule was formulated and to which manufacturing technologies and tools or materials it applies. The letters ‘AA’ describe the type of part to which the rule refers, i.e. ‘mold insert’ or molded and sintered ‘model’. Then follows the information about the production technique with a more detailed specification of tool group and material group. When a rule is applied to different production techniques or tool groups or material groups, the entries ‘xxx’ or ‘x’ are indicated. The rule ends with a consecutive number for the respective rule composition. Different rule sets exist for the mold insert manufacture and molding process of parts explained above. However, they can be geometrically connected. Hence

1.3 Design Rules

Fig. 1-2 Knowledge transformation by interpretation relevant to design

the structure details of the mold insert that are influenced by the manufacturing restrictions can also be found at the molded part, where the geometry sizes scale around the sinter shrinking and complementary structures are developed. The following section explains the design rules of the single technologies (especially of process preparation and production) that are connected to process chains for replication processes. A distinction is made between two large process chains, ‘micropowder injection molding’ and ‘microcasting’. 1.3.2 Design Rules Derived from Restrictions of Production Technology

Design rules are a methodical aid for achieving a knowledge transfer from technological facts (see, e.g., Fig. 1-5), especially from the operation system of production technology to the operation system of product development. This is demonstrated by the flow arrow in Fig. 1-4. Fig. 1-4 demonstrates that not only are there indications regarding manufacturing aspects passed on to the designer via guidelines for embodiment design and rules, as in mechanical engineering or precision engineering, but that it is also mandatory to employ the requirements and restrictions relevant to design that are included in the design rules. Without the active design that includes the knowledge facts in the rules, an effective and successful synthesis is not possible in primary-shaping microtechnology.

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Fig. 1-3 Classification scheme for design rules

The rules used at present refer exclusively to the production process chains for primary-shaped microcomponents. It has to be considered that most of the time the model is created in the CAD and from there a negative form – the mold insert – is generated. In the case of parametric CAD systems, the sinter shrinkage can also be considered in this process and the mold insert can be designed correspondingly larger. The manufacturing of the model is done in two steps. First, the mold insert (negative) is manufactured. The technologies available here are machining and abrasive processes (see Fig. 1-3 and Chapters 4–7). The second step is the molding process of the model (positive) by means of ceramic or metallic micropowder injection molding (Chapters 10–12) or microcasting (Chapter 13). To each of the two process steps different restrictions apply, e.g. it is not possible to mill a cavity that is smaller than the milling cutter. Also, micropowder injection molding requires a minimum wall thickness and a maximum flow length. However, both parts are geometrically unambiguously connected, i.e. restrictions of the mold insert manufacturing automatically apply also to the molded part and vice versa. Here it has to be considered especially that the geometric properties of both parts are not identical. First, the part is scaled during the molding process because of the sinter shrinkage and then a negative is cre-

1.3 Design Rules

Fig. 1-4 Flow of knowledge from production technology (PT) to product development (PD) by design rules

ated, so that, for example, a hole in the mold insert is converted into a cylinder with a decreased height and diameter.

1.3.2.1 Design Rules for Mold Insert Manufacturing Replication processes require as a first step the manufacture of a form – the mold insert. To achieve the aim of a cost-effective, medium-sized and largebatch production of microcomponents from metallic and ceramic material, abrasive and machining processes of the mechanical microproduction are more advantageous than processes based on lithography. For the majority of applications in the Collaborative Research Centre 499, micromilling has been used for the manufacturing of mold inserts so far. Among others, micro end mill cutters are employed here in order to manufacture 2.5- and three-dimensional microstructures. Process-specific parameters that have to be extracted are, for example, the body diameter of the tool and the length of the milling cutter’s edge that is linked to it. If interpreted with relevance to design, this implies that it is not possible to manufacture mold insert structures that are smaller than the milling cutter diameter plus the milling cutter tolerance or deeper than the maximum cutting depth. Owing to the circular cross section, vertical inner edges are also impossible, i.e. all mold insert edges – equivalent to outer edges of the final part – have to be provided with a minimal rounding radius.

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All these parameters and properties are collected in so-called technology specification sheets and are interpreted with relevance to design (Fig. 1-5) [21]. In systematic scenarios, the determined technological facts are now projected on to potential geometric structures or functional influences on the microparts. The results are descriptions and mathematical connections that correlate machine tool and tool parameters with design parameters. With the presentation of the design rules, the designer is provided with abstract and descriptive, but also concrete and computer-aided information about the same knowledge fact. Fig. 1-6 shows a design rule for three-axis micromilling that applies both to the end mill cutter and to the radius cutter. Various other design rules exist in addition that name the technological restrictions of three- and five-axis micromilling, and also rules for microelectrical discharge machining and for laser machining.

1.3.2.2 Design Rules for Replication Techniques The replication of microcomponents is done by micropowder injection molding (lPIM) and by microcasting. Micropowder injection molding as a replication technique for microcomponents differentiates between metallic and ceramic injection molding depending on the material to be molded. The lPIM process uses the mold inserts in order to mold the metallic or ceramic feedstock directly into these molds. The microcasting process is based on the lost-wax lost-mold technique, so as a first step models have to be manufactured. These lost models, mainly made from polymers, are mounted on a gate and feeding system made of wax. This assembly is completely embedded in a ceramic slurry. After drying, the ceramic is sintered, resulting in a ceramic mold with high mechanical strength. Simultaneously during the burning process, the polymer model is molten and burnt out. After the subsequent casting process, the metallic microcomponents can be taken out of the lost mold.

Micropowder injection molding Boundary conditions of the lPIM process result from the necessity to attach runners in a sufficient number and size to the part’s surface and to provide a surface for the ejector pins contacting the molded part for removal. The maximum achievable flow length and aspect ratios, and also sharp cross section transitions and cross section bendings, limit the mold filling behavior and the molding process quality. Especially the shrinkage of the material during sintering has to be considered. Therefore, taking into account the sintering shrinkage, it is possible that smaller structures may result compared with the dimensions of the mold insert, but shrinkage tolerances of ± 0.4% have to be considered at the same time (Fig. 1-7).

Fig. 1-5 Achievable manufacturing sizes in production preparation by means of three-axis micromilling [21]

1.3 Design Rules 15

Fig. 1-6 KR_FE_MF3_x_001 – Minimum radius for vertical inner edges

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Fig. 1-7 KR_MO_PIM_x_001 – Dimensional margin

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Microcasting To come to a design compatible with microcasting, several technological circumstances have to be taken into account. Compared with micropowder injection molding where the green compact and brown compact are intermediates on the way to the final sintered microcomponent, in microcasting more and versatile preparation and intermediate steps exist that influence the result. Depending on the manufacturing process for the lost models, different concepts for casting-compatible positioning of gates are necessary (model on substrate or single injection-molded models with gate and feeding system). The attainable surface roughness of the microcomponent is determined by the embedding mass employed and ranges down to Ra = 0.5 lm for Stabilor G. In addition, the attainable dimensional accuracy should be pointed out to the designer. By varying the expansion ratio of the embedding mass, the dimensional accuracy is adjustable within a few microns. In microcasting, small structures within a few 10 lm in wall thickness with at the same time high aspect ratios are processable. On falling below a specific structural diameter, the filling pressure rises in a hyperbolic manner, resulting in a more complicated form filling. This phenomenon is expressed by the design rule in Fig. 1-8, which consequently has an influence on the dimensional conception and the embodiment design (see Chapter 13) [22, 23].

1.4 Design Environment

The filing of rules in a database is important for the applicability of the rules for the designer and implementation in computer-based systems. The design rules can be provided over an interactive knowledge portal and/or directly with an application in 3D CAD. The design environment is planned in a way that supports the designer with respect to the product development phase in which the design is at that moment, by means of the design rules that are correspondingly altered with the abstraction level. In the early stages, the general comprehensive information about the production processes and the material properties are the interesting aspects. In the embodiment design phase, quantified values about realizable manufacturing sizes have to be provided for a detailed design draft. Here, concrete instructions are required that must be followed to realize a productioncompatible and functional part design. According to these different representation forms of the knowledge content of design rules, the design environment itself has to be planned to be flexible and application specific. Therefore, an information component, i.e. a web-based interactive knowledge portal, and an embodiment design component that is directly integrated into the CAD system as a working environment for the designer were developed for the product development of primary-shaped microcomponents. Both units access the same data of a database [24, 25].

Fig. 1-8 KR_MO_MIC_x_003: Minimum structure width

[22, 23]

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1.4.1 Information Unit

A ‘design and methodology database’ (KoMeth), which supports the designer as an interactive knowledge portal via different access possibilities with design rules, serves as an information component. The access page offers different selection fields that permit, according to the classification scheme (Fig. 1-3), the search for rules for a specific manufacturing technology, for specific part types or specific tools and materials. The rules that were found are displayed systematically in a hit list according to their rule class and name. If the designer is already familiar with the production process chain, a direct search over the list of existing rule classes and names is also possible. When selecting a rule from the hit list, an information page is displayed which shows the rule class, the numbering, the rule name, a prosaic description, an algorithmic–mathematical formula (if available) and sketches of geometric dimensions and illustrations of real manufactured microstructures (Fig. 1-9). In the course of product development, the abstraction level decreases and more and more technological restrictions have to be considered. Especially at the stage of realizing the embodiment design, the shape and dimensions of realizable structure details are not easy to comprehend any longer. In addition to the significance of errors and the inefficiency, these details deviate too much from the original design work. Therefore, the so-called design rule checker was established in mechanical engineering, microelectronics and mask-based microtechnology. This approach is also used for rule-based embodiment design of primary-shaped microcomponents. 1.4.2 Embodiment Design Unit

In addition to the presented qualitative and descriptive content, the knowledge base of the database also includes concrete values of single manufacturing technologies and knowledge about connecting these data with geometric characteristic values in the form of rules. These are formulated separately because both are independent of each other and subjected to permanent changes. The information and parameter values are derived from the state of the microproduction technologies and the material sciences. The rules are derived from the expertise and the experience in the design process, process preparation and production, and also from the part testing and material testing. Therefore, they are also subjected to permanent development and completion. All information and knowledge have to be saved independently of the CAD system. This offers the advantage that the same information and knowledge data can be used for different CAD systems and can be maintained without CAD.

1.4 Design Environment

Fig. 1-9 KoMeth – Design and methodology database

1.4.2.1 Preparing Elementary Rules for Computer-aided Design Rule Check With the application at hand, it has to be examined whether the designed geometry can be manufactured by means of the production technologies and materials selected. Therefore, the necessity to focus only on geometry is obvious, i.e. on the individual surfaces and edges. As a result, a description by means of boundary representation (B-rep) [26] has been chosen as the geometric model. In addition, many common transfer formats for CAD data such as STEP [27] or IGES support this description model.

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Fig. 1-10 Preparing elementary rules from general design rules and the geometry of the component

The task of a programmed algorithm is thus to derive geometric parameters from the three-dimensional geometric model and connect them by means of rules with the production technological parameter. By adapting general rules to the special geometry and to the individual production process, elementary rules are obtained. Consequently, they are not generally applicable any longer, but describe concrete circumstances and can hence be evaluated by the computer. As can be seen in Fig. 1-10, the preparation for this is carried out in two parallel paths: the adaptation and concretization of the rules and the determination of the geometric parameters.

Application- and process-specific adaptation of the design rules The processing of the rules as provided by a database until the time when they can be connected with the geometric parameters is carried out in three steps (Fig. 1-10, left path): 1. The rules are loaded from a database into a list corresponding to their rule code. Rules which are not appropriate according to the technology or the tool/material are not considered.

1.4 Design Environment

2. If a rule is defined for another type of machine element (model « mold insert), it is ‘translated’ by means of a transformation table. In doing so, certain key words are replaced by others (e.g. inner diameter ? outer diameter) 3. In accordance with the applied production process chain, the production technological data are loaded from a separate database, after which the respective placeholders in the rules are replaced. If no equivalent for a placeholder can be found in the database, the value ‘0’ is assumed.

Determination of the geometric parameters by means of the boundary representation method The evaluation of the geometry of the machine element is carried out in two phases (Fig. 1-10, right path). First, all boundary representations from the database of the CAD system need to be read, and second, the corresponding properties of each representation are to be determined. In the first phase: 1. a list of all solids of the machine element; 2. a list of all surfaces and edges of the solids from (1) are detected.

Geometric properties can be defined for solids, individual surfaces, two surfaces or edges (Fig. 1-11). For the determination of a property, geometric information about the individual boundary representations from the machine element database is available, such as unit (normal) vectors, fulcrum, limiting or proximate elements, etc. From this geometric information, simple geometric parameters can be derived. Two surfaces, for example, are parallel when they have identical unit normal vectors. For other parameters, the calculation is more complicated, since there are various influencing factors or interdependences. One example is the angle a of an edge, which is defined as the angle measured by the part between the adjacent surfaces directly at the edge. For the determination of this quantity, the fact is used that normal vectors are always indicated outwards from the volume solid of the CAD. In particular, the calculation runs through the following steps: 1. Determination of the adjacent surfaces. 2. Specifying a common point of the surface on the edge. 3. Determination of the unit normal vectors of two surfaces at this point. 4. Calculation of the angle between the unit normal vectors: cosu ˆ ~ n1
~ n2 :

Fig. 1-11 Determination of the boundary elements and their interrelations

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5. Determination of the small angle between the surfaces: a ˆ 1808 u : 6. Adjustment of nonparallel surfaces (08 < a < 1808), if it is an inner edge (a > 1808) or an outer edge (a < 1808). 7. Check if one of the surfaces exceeds the common edge in the direction of the other normal vector. This is possible by determining the turning point PE,1 on the surface F1 in the direction of the normal vector ~ n2 of the other surface F2. To obtain an obvious point, two further directions need to be defined. For this the cross product ~ n1  ~ n2 and the first normal vector ~ n1 are used. The point PE,2 is determined analogously. 8. Adjustment if one of these two points is on the common edge of the surfaces; if yes, then it is an outer edge, if not, it is an inner edge. 9. Calculation of the angle: with an outer edge: a ˆ a with an inner edge: a ˆ a ‡ 1808 : The detailed rules obtained by means of the steps listed above are now combined with all applicable boundary representations and their geometric properties in order to determine elementary rules. The latter consist only of numbers and mathematical operations and can be evaluated by an appropriate commercially available program [28, 29].

1.4.2.2 Design Rule Check The 3D CAD System Unigraphics for Knowledge Based Engineering (KBE) For realizing the knowledge-based design environment, commercially available system components are used to a great extent. Unigraphics (UG) V18 [30] is employed as a CAD system as it offers the possibility of a full parametric design. Furthermore, since version 17, a knowledge-based module (UG/Knowledge Fusion) (UG/KF) and a programmable user interface (UG/Open API) have been provided. The advantage of directly using a CAD system with an integrated knowledge-based module is that no neutral interface formats are required. As a result, costs and effort can be reduced and a loss of information (e.g. parametric or attached attributes) due to several format conversions can be averted.

Indication of the production process chain with associated technological specifications Before it is possible to check the microcomponent, the designer needs to indicate the production technologies used and the type of machine element of the product model that is to be designed. For this purpose, in a pop-up window a question on whether the mold insert or the model is involved and a question concerning the techniques and tools planned for the production are displayed. As a consequence, the corresponding rule classes can be derived.

1.4 Design Environment

Via the chosen rule classes, a database is addressed from which the applicable design rules on an abstract level are read out. In a further step, the parameters within the rules are replaced with concrete quantities from the production technologies, e.g. by the tool diameter or the process tolerances. A great advantage is that the production technological parameters are filed in a separate database. Especially in the microdomain they change permanently as there are always new process alternatives available or existing ones are improved. The filing is carried out separately for the production preparation and the production. In the production preparation not only the process but also the tools are distinguished. For each of these any properties can be defined, which are filed in a separate, connected database table. Hence new properties can be introduced via the database only by means of modifications. In a rule database all defined design rules are filed. They are contained in the form of an explanatory text, which is also displayed in case of a rule infringement, as well as in the form of an ‘IF – THEN – ELSE’ condition. The ‘IF’ part is formulated as a mathematical equation with placeholders for production technological and geometric parameters. If the described equation is correct, the ‘THEN’ part is carried out or, if not, the ‘ELSE’ part, which may be e.g. an automatic correction.

Design Rule Check (DRC) process The checking is to a great extent carried out by means of a C++ program, which has been written exclusively for this task. In this program, the CAD data and databases are read in and elementary rules are created. These elementary rules are checked in the UG/KF module and rule infringements are presented in a pop-up window. In order to demonstrate functionality, a micromechanical component has been defined, with the aid of which many production restrictions can be presented. It is a short T-piece at which the web has been shortened slightly and the vertical edges are rounded so that the mold insert can be manufactured by micromilling. The mold insert is to be milled with an end mill cutter (diameter 200 lm, length of cutting edge 400 lm) on a three-axis micromilling unit. To induce a rule infringement, the edges of the web are not rounded. After the adjustment of the production technology and the type of machine element, the checking is started and the result is later presented in a dialogue box (Fig. 1-12 a). When an infringed rule is selected, the matching geometric element is marked in the CAD model and additional information concerning the rule is displayed. In this case (not rounded edge) also an automatic correction is possible. If this is carried out for all four edges, the program does not display an error message any longer (Fig. 1-12 b). If the corrections are acceptable for the designer, drawings for the mold insert and the model can be generated for production and quality assurance. Alternatively, it might be possible to transfer the design data directly via a CAD–CAM interface to the production preparation.

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a)

b) Fig. 1-12 (a) Interactive window with infringed rules; (b) manufacturable part after automatic correction (rounded vertical edges)

1.6 References

1.5 Conclusion

The approach presented here for the design of primary-shaped microcomponents and microsystems was defined and validated by developing and designing a microplanetary gear. Therefore, a tolerance concept for generic and gearing tolerances was established and implemented in the gear unit. By manufacturing and measuring single components of the gear, perceptions are iteratively integrated to the tolerance concept. Subsequently, our Institute will introduce a microgear test rig in order to test the components of microgears. The goal is to obtain information regarding the transmission behavior of high-strength microcomponents that are in contact within the system. In combination with the dimensional measurement of single components, microsensitive features and properties of microgears are deduced with consideration of their effects on function. Moreover, the development of a new VDI guideline is being directed by the Institute owing to its activities within the domain of microgears.

1.6 References 1 G. K. Fedder, ‘Structured design of inte-

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grated MEMS’, in: IEEE Twelfth International Conference on MicroElectro Mechanical Systems; 1999, pp. 1–8. K. Hahn, Methoden und Werkzeuge zur fertigungsnahen Entwurfsverifikation in der Mikrotechnik; Dissertation, VDI-Verlag, Düsseldorf, 1999. M. Kasper, Mikrosystementwurf; Berlin: Springer, 2000. M. K. Long, J. W. Burdick, E. K. Antonsson, ‘Design of compensation structures for anisotropic etching’, in: MSM 99; 1999, pp. 124–127. St. Büttgenbach, U. Hansen, L. Steffensen, Konstruieren in der Mikromechanik; Düsseldorf, VDI-Verlag, 2000. J. Mehner, Entwurf in der Mikrosystemtechnik; Habil., Dresden University Press, Dresden, 2000. J. M. Rabaey, A. Chandrakasan, B. Nikolic, Digital Integrated Circuits – A Design Perspective, 2nd edn.; Upper Saddle River: Prentice Hall, 2003. W. Menz, J. Mohr, O. Paul, Microsystem Technology; Weinheim: Wiley-VCH, 2001.

9 Ch. Lessmöllmann, Fertigungsgerechte Ge-

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staltung von Mikrostrukturen für die LIGATechnik; Dissertation, Karlsruhe, 1992. P. Buchberger, MIDAS Ein wissensbasiertes System zur Unterstützung des fertigungsgerechten Entwurfs von LIGA-Mikrostrukturen, Bremen; 1997. B. Forner, R. Feiertag, Konstruktionsunterstützung für LIGA-Mikrostrukturen, BMBF-Verbundprojekt METEOR; Wuppertal, 1997. W. Ehrfeld (ed.), Handbuch Mikrotechnik; München: Hanser, 2002. G. Pahl, W. Beitz, Konstruktionslehre, 5th edn.; Berlin: Springer, 2003. G. Pahl, W. Beitz, Engineering Design, 2nd edn.; London: Springer, 1996. VDI-Richtlinie 2221, Systematic Approach to the Development and Design of Technical Systems and Products; Berlin: Beuth, 1993. VDI-Richtlinie 2223, Systematic Embodiment Design of Technical Products; Berlin: Beuth, 2004. K. D. Müller-Glaser, ‘Moderner Entwurf von Mikrosystemen’, e&i, 1997, 114, 151– 159.

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thodisch unterstützter Entwurf mechanischer Mikrosysteme, ETG-/GMM-Fachtagung: Innovative Klein- und Mikroantriebstechnik; Berlin: VDE Verlag, 2004, pp. 63–67. A. Albers, J. Marz, ‘Restrictions of production engineering on microspecific product development’, Microsyst. Technol. 2004, 10, 205–210. J. Marz, N. Burkardt, A. Albers, ‘Methodological investigation of the product development in microtechnology,’ in: Proceedings of the 4th International Topical Conference, European Society for the Precision Engineering and Nanotechnologies (EUSPEN); 2003, pp. 23–26. A. Albers, J. A. Marz, N. Burkardt, in: Design Methodology in MicroTechnology, 14th International Conference on Engineering Design, ICED 03; 2003. G. Baumeister, N. Holstein, V. Piotter, R. Ruprecht, G. Schanz, M. Guttmann, F. Winkler, A. Häfele, ‘Herstellung metallischer Mikrobauteile unter Einsatz verlorener Kunststoffformen’, Nachrichten FZK 2002, 34, 198–209. G. Baumeister, K. Mueller, R. Ruprecht, J. Hausselt, ‘Production of metallic high aspect ratio microstructures by microcasting’, Microsyst. Technol. 2002, 8, 105– 108. A. Albers, N. Burkardt, S. Hauser, J. Marz, ,Prototyp einer wissensbasierten

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Konstruktionsumgebung für den Entwurf von Mikrobauteilen‘ Konstruktion 2005, 57, 76–81. A. Albers, N. Burkardt, S. Hauser, J. Marz, ‘Knowledge-based design environment for primary shaped microparts’, Microsyst. Technol. 2004, in press. R. P. Cherian, P. S. Midha, L. N. Smith, A. G. Pipe, ‘Knowledge based and adaptive computational techniques for concurrent design of powder metallurgy parts’, Adv. Eng. Soft. 2001, 32, 455–465. Geometric and Topological Representation; DIN V EN V ISO 10303-42 (1994), 1994. Maplesoft, Maple 9.5, http://www.maplesoft.com/products/maple, September 2004. The MathWorks, Matlab 7, http:// www.mathworks.com/products/matlab/ ?BB=1, September 2004. Unigraphics, CAD/CAM/CAE-System für die durchgängige virtuelle Produktentwicklung (VPD), http://www.ugsolutions.de/ products/unigraphics.shtml, 2002. V. Piotter, T. Gietzelt, K. Müller, R. Ruprecht, ‘Herstellung von metallischen und keramischen Bauteilen durch Pulverspritzgießen’, in: Material- und Verfahrensentwicklung für mikrotechnische Hochleistungsbauteile; 2001, pp. 59–64. M. Weck (ed.), Mikromechanische Produktionstechnik, DFG SPP 1012; Aachen: Shaker, 2003.

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2 Modeling and Validation in Design A. Albers, D. Metz, Institute of Production Development (IPEK), University of Karlsruhe (TH), Germany Abstract Microspecimens and microcomponents were modeled in order to analyze the influence of the material anisotropy of the grains upon the stresses which develop during loading. The limit of finite element analysis with isotropic material models and the necessary consideration of the microstructure was investigated. The influence of the number of grains in critical areas on the local stress was analyzed. This has to be taken into account for a reliable dimensioning of microcomponents. In a further step, the optimization of microcomponents was attempted by means of an optimality criteria method, in order to minimize the stresses at critical locations.

Keywords finite element analysis; stress distribution; microstructure; grain orientation; elastic anisotropy

2.1 2.2 2.2.1 2.2.2 2.2.3 2.2.4 2.2.5 2.2.6 2.3 2.3.1 2.3.1.1 2.3.1.2 2.3.1.3 2.3.2

Introduction 30 Modeling 30

Modeling of the Grain Structure 31 Modeling of the Microstructure 31 Modeling of the Single Grain Material Mesh Generation 33 Microspecimens 34 Microcomponents 35

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Simulation Results 35

Simulation of Microspecimens 36 Microbending Specimens 36 Microtensile Specimens 37 Microtensile Specimens Containing Pores 39 Microcomponents 42

Advanced Micro and Nanosystems Vol. 3. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31208-0

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2.4 2.4.1 2.4.2 2.4.3 2.5 2.6 2.7 2.8

Structural Optimization of Microcomponents

44

Basic Principles of Shape Optimization 44 Modeling of the Shape Optimization 45 Results of the Shape Optimization 45 Conclusion 48 Outlook 49 Acknowledgment References 49

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2.1 Introduction

To reduce development times and development costs, a modern product development process increasingly often employs numerical methods in order to replace costly tests or at least to reduce their number. By means of the finite element method and modern structure optimization tools, components and systems are produced and tested virtually. Stresses are determined, dimensioning is employed and variants are compared. In this way, long and expensive iterative development processes can be shortened and the number of prototypes can be reduced. Especially in the field of microtechnology, time- and cost-saving numerical processes should be used owing to the major expense of the production of prototypes. However, as in the dimensions of microtechnology effects occur that can otherwise be neglected, the question is raised of whether common numerical methods can be applied without any modifications. Therefore, this chapter deals with the question of the reliable simulation of microcomponents and the reliable transfer of macroscopic modeling to the micro range. Based on these simulations, microcomponents and microsystems are optimized with regard to the stresses that occur and reliable data for the dimensioning of microcomponents are presented for the design engineer.

2.2 Modeling

In the Collaborative Research Center SFB 499, tests are carried out with specimens (for microtensile and micro three-point bending tests, see Chapters 19 and 20) and with microcomponents and microsystems (see Chapter 1). In addition to the experiments, numerical analyses are carried out in order to examine the specific influences of the microstructure and to permit the dimensioning of highly stressed microcomponents during design. Here it is necessary to model the microspecimens and microcomponents with regard to their dimensions and their ‘inner structure’, i.e. the grain structure.

2.2 Modeling

2.2.1 Modeling of the Grain Structure

The microcomponents and the microspecimens are produced by powder injection molding using the materials zirconia and 17-4PH (sintered steel) or by microcasting of Stabilor G (gold-based dental alloy). The characteristic dimensions of the examined components vary from 0.2 to 1 mm. The characteristic grain size of the materials employed varies from about 0.5 to 1 lm for ceramics and from 20 to 50 lm for metals. The finite element method usually assumes isotropic material behavior. However, anisotropic material behavior can also be taken into account when simulating single crystals or compound materials. Apart from these and some other cases, almost always quasi-isotropic material behavior can be assumed, as numerous grains exist in the workpiece. The assumption of isotropic material behavior is no longer justified when dealing with microcomponents that have only a few grains in the relevant cross-section. In this case criteria must be established for the relationship between microcomponent size and microstructural parameters such as the grain size which allow the consideration of anisotropic grain behavior [1, 2]. In order to investigate these questions, a model for polycrystalline microstructures was developed. 2.2.2 Modeling of the Microstructure

For the modeling of polycrystals, some basic assumptions are applied: · The grains are large enough to describe them within the scope of continuum mechanics. · The material behavior of the grains is described with a linear-elastic material model. · The grains are assumed to be defect free, hence the material behavior is described with the ideal single crystal parameters. (Since no single crystal data are available for Stabilor G, they are approximated by single crystal data for gold.) Stochastic geometries are used for the modeling of irregular polycrystals. Owing to its statistical properties, the Voronoi tessellation is suited to describe the geometry of a polycrystal [3, 4]. The Voronoi tessellation is often used for the generation of random grain structures, especially in material research. A given number of points is taken as a starting point of the Voronoi tessellation. These so-called generating points result from a process for describing a random tessellation of a region [5, 6]. This point-generating process is chosen in such a way that enables the Voronoi tessellation to describe the microstructure of the polycrystals as realistically as possible, possessing also simple statistic properties. A Voronoi tessellation is generally understood as a tessellation of a region into convex polyhedrons, which do not overlap and consume the whole

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Fig. 2-1 Micrograph of a microspecimen of zirconia (see Chapter 20) and Voronoi tessellation for mathematical description

area. Fig. 2-1 shows a graph of a Voronoi tessellation. Owing to the properties of a polycrystal, it is reasonable to use homogeneous and isotropic point processes, such as the Poisson process [3, 5, 6]. Voronoi tessellations, which are generated by a Hardcore Poisson process, are suitable for describing the grain structure of a regular polycrystal [2, 3, 7]. In order to model the grain structure of the materials as accurately as possible, micrographs are compared with the numerically generated models (see Fig. 2-1). A good correspondence was reached by variation of the Hardcore distance. 2.2.3 Modeling of the Single Grain Material

Two assumptions are made when modeling the three-dimensional material behavior of a polycrystal: · The polycrystal is macroscopically isotropic, i.e. a preferred direction (texture) of the crystallographic axes of the single grains does not exist. · The orientation of neighboring grains is interdependent. The orientation of a crystal is usually described with the three Euler angles W, H, and U with 0 £ W < 2 p, 0 £ H < p and 0 £ U < 2 p in crystallography [8]. The global coordinate system of the model is defined in a first step. The local coordinate system is connected to the global coordinate system via rotations around the three Euler angles. The three-dimensional material behavior of the single crystal is given by the stiffness tensor regarding the local coordinate system. A random orientation based on single crystal values is allocated to each grain. Here, the Euler angles are used as random variables. Then the orientation of each single grain is a random combination of these variables. An infinite number of grains result in an isotropic polycrystal. In order to describe an al-

2.2 Modeling

most isotropic material, the Voronoi tessellation has to provide a sufficient number of grains [1, 2]. Different material property files are used for the anisotropic FEM calculations and their further statistical evaluation. Each material property file contains material data for the grains and is based on the results of single crystal testing combined with the random three-dimensional orientation of these crystals. 2.2.4 Mesh Generation

Owing to the large stress gradients that have to be expected owing to the discontinuity in the material behavior of neighboring grains, it is necessary to mesh the grains in a sufficiently accurate way. Based on the algorithm VorTess [9] and the works of Weyer et al. [10], a rectangular, two-dimensional area with a random grain structure is generated by means of a C program. Since an intact grain structure is assumed, the meshes of neighboring grains have to be compatible with each other. The meshing of the Voronoi tessellations is done in MSC.Patran, because the Patran command language [11] provides an efficient macro language that is suitable to mesh higher numbered grain models in a compatible manner within finite time. The size of the finite elements is given by the globular structure of the Voronoi tessellations. The largest value is determined by the smallest edge of two neighboring grains in the whole model, because at least one finite element has to describe this area of the grain. Owing to the large fluctuations in the length of the grain boundaries, the smallest edge often represents a reasonable value that should only be replaced by a lower value in the case of models with extremely few grains. For more than 25 grains, a finite element modeled microspecimen or microcomponent with grain structure shows an essentially higher number of finite elements than a typical model without represented grain structure. In the finite element method, the number of finite elements has a significant influence not only on the quality of the calculation, but also on the calculation time. Owing to the large number of calculations necessary to deliver a reliable statistical result, it is desirable to limit the number of finite elements to the least necessary. Therefore, a different meshing strategy was applied for single grains in order to reduce considerably the calculation time for the models with grain numbers above 200. In the area of the grain boundaries, the size of the finite elements in the model was determined by the smallest edge. However, the size of the finite elements in the interior of the grains was substantially increased and therefore the number of finite elements – and as a consequence the calculation time – could be considerably reduced. Nevertheless, the accuracy of the calculation could be maintained.

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2.2.5 Microspecimens

Simulation of grain structures usually employs the concept of a representative volume element [12, 13]. Two different approaches on the micro and on the macro levels are applied for describing the material behavior [14–16]. In doing this, the following requirements have to be met: First, the micro level must be sufficiently large, i.e. the considered area must provide a sufficient number of microstructural details and can therefore be considered to be representative. Second, the considered section must be sufficiently small that the stress gradients existing on the macro scale are negligible and the material on the macro level can be considered as homogeneous. However, the dimensions of the microspecimens or microcomponents that are treated in the Collaborative Research Center with their respective grain sizes and also the stress gradients that occur do not always comply with the two requirements described above. As for microcomponents or microspecimens of ceramics (zirconia), sufficient numbers of grains in critical cross-sections are even provided in the ‘worst case’, i.e. the smallest part dimension 160 ´ 160 lm and largest grain size 0.3–1 lm. Here, (quasi-)isotropic material behavior can be assumed. This is not the case for metals (17-4 PH, Stabilor G) with a grain size of 20–50 lm. In order to guarantee reliable modeling and simulation of the microspecimens or microcomponents, the microstructure has to be modeled either directly in the microspecimens and in the microcomponents or in the region of the component which is of major interest. Owing to an appropriate choice, the geometry of the Voronoi tessellations could be directly used for the modeling of the microspecimens as microspecimens. The upper part of Fig. 2-2 shows a micrograph of a microbending specimen and the lower part shows a modeled microspecimen with 200 grains discretized with 176 000 finite elements, which are not visible here. By considering

Fig. 2-2 Micrograph of a metallic microbending specimen with 200 lm width (see Chapter 19) and modeled microspecimen with 200 lm width in MSC.Patran with 200 grains

2.3 Simulation Results

the results given in Chapters 19 and 20, it was possible to model the mechanical behavior of microspecimens two-dimensionally with regard to the grain distribution. The comparison of micrographs of a metallic microbending or tensile specimens showed good correspondence of the mathematical models with regard to the specimen’s grain structures. 2.2.6 Microcomponents

The influence of the grain anisotropy in the case of complex geometries was studied by the example of the ratio of critical cross-sections to the grain diameters of sintered metals similar to those in the microspecimens. As shown in Fig. 2-3, the tooth of a microplanetary wheel (see Chapter 1) was memorized into a MSC.Patran database in which a Voronoi tessellation had been generated previously. The contour of the part serves as a boundary. Unnecessary finite elements of truncated and entire grains outside of the profile were removed. Then the nodes were positioned on the circumferential edge of the model in order to describe the profile exactly.

2.3 Simulation Results

The aims of the following calculations are to carry out fundamental examinations of the influences of the grain anisotropy and to find the limits of the isotropic modeling of different materials.

Fig. 2-3 Tooth of a microplanetary wheel with 33 grains and meshed grain structure

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Because of the small grain diameters of zirconia, from 0.3 to 1 lm, and the characteristic geometric dimensions of the microspecimens and microcomponents, from 130 to 260 lm, zirconia can be considered as quasi-isotropic. Since zirconia shows a distinctive elastic anisotropy [16] and further miniaturization is intended, the influences of the grain anisotropy and the limits of the isotropic modeling are still to be analzyed. Therefore, the considered geometric dimensions have to be decreased for zirconia compared with the models for metallic materials. The specimen models with 10 grains in the cross-section have a 200lm width for metallic materials and only a 10-lm width for zirconia. The same applies for the microcomponents. In the following, the results always refer to the number of grains in critical cross-sections. It should be kept in mind that this means different dimensions for ceramic and metallic microspecimens or microcomponents. To establish the influences of the grain anisotropy, the anisotropic results are standardized with the isotropic results. HKS/Abaqus was used as finite element solver. 2.3.1 Simulation of Microspecimens

Micro three-point bending tests and microtensile tests were employed for the determination of material values (see Chapters 19 and 20).

2.3.1.1 Microbending Specimens Models for 10 different microbending specimens were generated two-dimensionally, with a width to length ratio of 1 : 5. The total number of grains ranged from 10 to 250 and the number of grains in the critical cross-section from 1 to 9. The microbending simulation was carried out linear elastically with contact and the seating and the die were modeled as rigid bodies. The travel of the die was about 4% with respect to the width. The models were calculated for plane stress with the materials zirconia, 17-4 PH and gold. The models with 5–9 grains in the width of the specimens describes the grain number for the metallic specimens with 200 lm width. In order to examine the limits of conventional modeling and further simulation, each model was first calculated isotropically, i.e. without considering the grain structure. Then the single models were calculated with 50 different anisotropic material property files, i.e. considering the grains. For the evaluation and comparison of the results, the strain energy of each model was analyzed and standardized with the isotropic result. Fig. 2-4 shows the deviations of the standardized strain energy for microbending tests with 17-4 PH. Qualitatively similar results were obtained for zirconia and gold. At a nearly constant average value, the scatter of the strain energy decreases with increasing number of grains, and the standard deviation thereby drops continuously from 5.5 to 1.4%. The same analyses were carried out with zirconia and gold. Here the results are qualitatively similar, but the deviation of the

2.3 Simulation Results

Fig. 2-4 Deviation of the strain energy in the microbending simulation with a specimen made of 17-4 PH

strain energy in the case of zirconia is twice as large owing to the distinctive material anisotropy. In the case of 17-4 PH and gold, it converges from initially a 10% deviation to 2%, and in the case of zirconia from 20 to 10%. By analyzing the stresses in microbending simulations a relatively large scatter could be observed at the tensile side in the middle of the microspecimens of up to 30% with zirconia and 22% for 17-4 PH and gold, in comparison with the isotropic calculation. The increased stress always occurred when a grain with an ‘unfavorable’ orientation was in this area. If only a few grains are located in the bearing cross-section, the dimensioning process has to consider this case. The data lead to the estimation that for more than eight grains in the relevant crosssection, quasi-isotropic conditions can be assumed for the metals considered. Owing to the bending load stress gradients occur in the models. Therefore, the influences of the grain anisotropy on the stresses of critical cross-sections or critical areas cannot be described exactly.

2.3.1.2 Microtensile Specimens Tensile tests were simulated for metals and for ceramics in order to elucidate the influences of the grain anisotropy on the stresses. The analysis of the microtensile test was similar to that of the microbending test. Ten different micromodels of tensile specimens were generated two-dimensionally with a width to length ratio of 1 : 5. The number of grains ranged from 10 to 250 and the number of grains in the critical cross-section from 1 to 8. The microtensile test was modeled linear elastically. The nodes at the lower end of the specimen were kept at a fixed position and the nodes at the upper end were charged with a force in the tensile direction. The models were calculated

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Fig. 2-5 Left, homogeneous von Mises stress distribution in the case of isotropic calculation; middle and right, inhomogeneous von Mises stress distribution in the case of different anisotropic calculations

for plane stress. Similarly to the bending models, a number of grains from 5 to 8 in the width of the specimens describes the metallic specimens with 200 lm width. In order to examine the limits of the conventional modeling, each model was first calculated isotropically, i.e. without considering the grain structure. Then the single models were calculated anisotropically with 50 different material property files. Fig. 2-5 shows a modeled microtensile specimen with 70 grains, on the left the isotropic calculation with a homogeneous state of stress, and in the middle and on the right three anisotropic calculations. Fig. 2-5 illustrates the inhomogeneous stress state due to grain anisotropy. The stresses of the single grains show a clear dependence on their orientation. For each calculation, another stress distribution can be observed in the model. The calculations with different numbers of grains show similar results. Furthermore, two different grain structures with 50 different anisotropic material property files were calculated for each number of grains. Both models show very similar results. The results are presented in Fig. 2-6, which shows the scatter of the strain energy of 100 anisotropic calculations for the models with different grain numbers with respect to the particular isotropic case. It can be observed that the scatter of the strain energy decreases with increasing number of grains in the models. A reduction in the deviation from the average value is expected and would approach the isotropic limit for an infinite

2.3 Simulation Results

Fig. 2-6 Scatter of strain energy with respect to the isotropic value

number of grains in the model. The standard deviation of the strain energy drops continuously with increasing number of grains from 9 to 1%. The scatter of stresses and the strain energy in the case of zirconia is larger than for 17-4 PH or gold, but from a qualitative point of view similar results can be assumed for these materials. Here the results allow the estimation that for more than eight grains in the relevant cross-section, quasi-isotropic conditions can be assumed.

2.3.1.3 Microtensile Specimens Containing Pores Material tests (see Chapter 19) showed pores as a possible failure initiation area in the case of tensile loading. Therefore, the influence of the grain structure on the stresses at pores was analyzed using further models, each with one pore at a different position, set up on the basis of an existing two-dimensional model for microtensile tests with 210 grains. The pore size was nearly half the diameter of a grain, so that the pore always influenced several grains. Gold was used as material, because pores sometimes occurred in Stabilor G microspecimens made by microcasting. The models of the microtensile test pieces containing pores were calculated with the same material property files as the respective models without pores. Then their results were compared. The resulting strain energies did not show significant differences for models with or without a pore, because the influence of the pore was very local only. However, in the area around the pore a noticeable stress increase vertical to the tensile direction occurs (see Fig. 2-7). In Fig. 2-7 the four locations where the highest stresses occured are labeled. In order to examine the influence of the pore on the local stress, the respective stress maxima were compared with the

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Fig. 2-7 Position of the pores and simulation results, each for an anisotropic calculation

stress values at the same position for the same model without pores. Owing to the grain structure, it was not possible to determine a nominal stress value for the cross-section without pores. Therefore, the stress increase can be understood only in a limited manner as a notch factor. Fig. 2-8 shows the scatter of the stress values at four locations of the anisotropic calculation with respect to the stress value at the same locations without pores. The two fluctuations on the left side apply to the left and right areas of the pore on the edge; the deviations on the right side apply to the left and right areas of the pore in the middle of the microtensile specimen. The two-dimensional modeling of a pore can be interpreted like a hole in a flat bar for an analytical calculation. The notch factor for an infinitely thick flat bar is then 3 [17]. The scatter shown in Fig. 2-8 is decisively influenced by the orientation and size of the grains that directly surround the pore; the other grains only have a small influence. The orientations of the individual grains in the area of the pore were varied specifically in order to explain the deviations of the stress increase. One materi-

2.3 Simulation Results Fig. 2-8 Stress scatter at a pore

al property file with a high increase factor and one with a low increase factor for locations 1 and 3 were further examined. The orientation of the grains in the area of the pore with the higher increase factor (here called material a) was varied by the orientation of the same grains of the lower increase factor (here called material b). Table 2-1 shows in the upper row the increase factor at location 3 of the model with the pore in the middle of the microtensile specimen. The lower row displays the reduction of the increase factor at location 1 of the pore at the edge of the microtensile specimen. Table 2-1 clearly shows the approximation of the stress increase of material a on material b due to the variation. The stress increase is significantly influenced by the orientation of the adjacent grains. The remaining difference is caused by the influence of the other neighboring grains. If they are also changed, the value of material b is almost obtained. The current simulations correlate well with the theoretical factor 3 of a cylindrical through-thickness pore. However, no findings exist about the effects in the direction perpendicular to the model. Therefore, further three-dimensional simulations are necessary. Until this is implemented, a slight safety addition to the dimensioning according to [17] should be employed.

Table 2-1 Reduction of the stress increase factor due to variation of the grain orientations

Location 3 Location 1

Material a

Variation: upper grain

Variation: both grains

Material b

3.16 3.22

3.06 3.16

2.84 3.05

2.65 2.8

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2.3.2 Microcomponents

Results obtained from microspecimens with simple geometries cannot be transferred without further considerations to real components with a more complex loading. For this reason, the influence of the grain structure on actual microcomponents was also examined. Five different models of a single tooth of a microplanetary wheel were generated for the examinations. The grain number in the tooth varies from 17 to 97. The single models provide 3–7 grains in the cross-section of the tooth root. The model with 97 grains represents an upper limit. The modeling of a larger number of grains in the tooth of the microplanetary wheel fails owing to the extent of the modeling, which was done manually until now. The calculations were performed with the materials 17-4 PH, zirconia and gold. The models were calculated assuming plane stress. Similarly as in the models of tensile and bending specimens, a number of grains between five and seven describes the structure of gearwheels made of metals in the tooth root cross-section. In order to reduce the calculation time, a direct force on the tooth was used instead of a contact model. The lower edge of the models was fixed. The examinations were standardized with the results of the isotropic case again. The left side of Fig. 2-9 shows the isotropic calculation with a direct flux through the tooth profile. In the middle an anisotropic calculation with a stress maximum at the right tooth root is displayed. On the right an anisotropic calculation with a decisive stress increase in the left tooth root is shown. The models show similar results for all numbers of grains. In both anisotropic calculations a grain with a low stiffness towards the force flux can be observed close to the load application. However, since no critical location exists, this can be neglected. The stress maxima in the anisotropic case preferably occur at the left tooth root,

Fig. 2-9 Calculation of a tooth of a microplanetary gearwheel with a total of 33 grains and with five grains in the tooth root cross-section for zirconia: on the left the isotropic calculation, and in the middle and on the right the anisotropic calculations

2.3 Simulation Results

but in one-third of the calculated cases, it occurred at the right tooth root also. The stress maxima occurred in all areas of the tooth root, but preferably at grain boundaries, as shown at the right tooth in Fig. 2-9. This can be explained by different stiffnesses, depending on the orientation of the adjacent grains. The results often show an almost direct flux of the stress through the tooth, similarly to the isotropic calculation (see Fig. 2-9). In the case of anisotropic material behavior, deviations of the stress values in the tooth root are observed for different material property files. These deviations and the behavior described above could be observed for all materials examined (17-4 PH, zirconia and gold). Solely the grain or grains directly located in the tooth root decisively influence the stress value in this region. This could be demonstrated by specific variation of the orientations of single grains in the tooth root. All materials show qualitatively the same behavior, with the greatest scatter for zirconia and the smallest for gold. Owing to the dependence of the stress values on the orientation of single grains, it was not possible to obtain a convergence of the scatter of the stresses relative to the isotropic case by increasing the grain number at the left tooth root. The standard deviation was 20% for zirconia, 12% for 17-4 PH and 7% for gold. The right tooth root showed a slight convergence of the standard deviation from 20% to 16% for zirconia, from 16 to 12% for 17-4 PH and from 11 to 7% for gold. In the worst case, anisotropic calculations showed a large stress increase of up to 80% with regard to the isotropic calculation. Owing to these extreme stress increases, a factor of 1.8 has to be included with respect to the safe dimensioning for microcomponents of zirconia with only a few grains in the cross-section. For 17-4 PH and gold a factor of 1.4 is sufficient. The variation of the orientation of single grains, corresponding to those of the calculations of microtensile tests with pores, also showed the great relevance of the orientation of the grain at the critical position. It has to be added that, owing to the two-dimensional modeling, the influence of further grains in the third dimension could not be considered, so that only one or two grains influenced the deviations. However, there are more grains across the tooth width in the highly stressed area, so that by an averaging over the grains in the tooth width, a reduction of the deviation and therefore a reduction of the mentioned factors of 1.8 and 1.4, respectively, is probable. This is why a three-dimensionally modeled grain structure of microplanetary wheels is examined. It is desirable to reduce the large stresses in the tooth root already during the design process. For this purpose, mechanical engineering employs numerical optimization processes in order to decrease the stress at highly stressed areas in workpieces by modifying the shape.

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2.4 Structural Optimization of Microcomponents

The examinations of microgearwheels clearly show the large stresses at the most critical position for failure at the tooth root. If it were possible to reduce this stress, a larger torque could be transmitted. For the optimization of components, mechanical engineers employ optimality criteria methods, such as shape optimization. The question is whether this method can be applied to microcomponents. 2.4.1 Basic Principles of Shape Optimization

The optimality criteria methods take advantage of the knowledge of the physics and mechanics of the respective problem set. Theses will be postulated describing the optimum. A well-known and ascertained physical law relating to structural mechanics is the fully stressed design [18]. An important mathematical optimality criterion is the Karush–Kuhn–Tucker condition [19, 20], normally designed for convex optimization purposes. The theses of stress homogenization and stress minimization can also be considered as optimality criteria [21]. Regarding the optimality criteria methods, these criteria and the response behavior of modifications of the physical model are implemented into a suitable algorithm. A convergence behavior is achieved by the use of suitable redesign rules. The optimality criteria are particularly well proven for shape and topology optimization where a large number of design variables is required. It is important to note that the convergence rate is independent of the number of design variables. The optimization program used, TOSCA.shape, implements an optimality criteria approach. A distribution of scalar values within a design area, the so-called design response, are basically processed by the optimization algorithm. Depending on the user-defined objective function, these values are minimized or maximized by means of a homogenization within the design area. The shape change induced by the optimizer, i.e. the variation of the finite element mesh, is managed by displacing nodes. This leads to a decrease in the local surface curvature in highly stressed regions. Typically, such scalar quantities are von Mises stresses obtained from a static finite element analysis. The nodal displacement DU that is applied to a certain node by the optimizer is then calculated by means of the following equation: DUn ˆ aF…a; b; . . .†

…1†

where a denotes a scalar calibration factor and F(a, b, . . .) is a function of the scalar nodal design response a, used for optimization such as von Mises stress, a reference value b and other parameters.

2.4 Structural Optimization of Microcomponents

2.4.2 Modeling of the Shape Optimization

The finite element calculation of the microgearwheels is followed by a shape optimization. A tooth of the microgear with two grains in the left tooth root was chosen as a model for optimization. The surface nodes in the left tooth root were determined as design area and as design response and their von Mises stress values were determined. The objective function was stated as: minimize the stress distribution within the design area. Two boundary conditions were included in order to achieve reasonable results in the calculation and with respect to the running ability of the gear and the feasibility of production [22]. A mesh smoothing, i.e. the retightening of the surface node that borders the design area and the prevention of shrinking of the design area in order to achieve a continuous transition from design area to tooth flank and to tooth base, respectively, was set up. The iterative process ends when a user-defined stop condition is fulfilled. In this case, the optimization was aborted after 15 iterations, when no substantial difference could be determined any longer. 2.4.3 Results of the Shape Optimization

First, the model was optimized without consideration of the grain structure and with isotropic material data. Then the same model was calculated and optimized with 25 different anisotropic material property files. The left side of Fig. 2-10 shows the different contours of the notch root for the isotropic material and three anisotropic material property files (here denoted materials c, d and e). The orientations of the individual grains in the root of the tooth were varied specifically in order to explain the different contours after optimization. The orientations of the grains in the tooth root from material c were varied by the orientation of the same grains of material e. On the right side, the approximation of the tooth root contour in the case of the variation of the grains’ orientation in the tooth root is shown. As usual for shape optimization, small changes in the contour have a distinctive effect on the stresses (see Figs. 2-10 and 2-11). The results of the anisotropic material property files differ noticeably from the result for the isotropic material. However, compared with the original models, a reduction in the tooth root stress from 10 to 32% could be achieved for all 25 optimizations. In order to demonstrate the influence of the orientation of the single grains, the individual grain orientations in the tooth root were varied specifically. Material c was varied by the orientation of the same grains of material e. For this purpose, the orientation of the upper grain in the tooth root in material c was replaced by the orientation of the grain of material e. Then, additionally, the orientation of the upper grain was replaced. The difference in the geometry con-

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Fig. 2-10 Left, tooth contours after the optimization with different material property files and right, convergence of the tooth contour for the variation of the grain orientation and renewed optimization

Fig. 2-11 Proportional deviation of the tooth root stress at the optimized geometry with respect to to the non-optimized tooth root

2.4 Structural Optimization of Microcomponents

tours of the varied model can be explained by the influence of the remaining surrounding grains (see Fig. 2-10). After having clearly demonstrated the influence of grain anisotropy, the question is how the geometry change affects the tooth root stress when other material property files are used for calculation with the new geometry instead of those used at optimization. This is relevant because in practice a large number of components with very different grain structures will be molded by means of one mold with a geometry defined after optimization. Therefore, the possibility of optimizing the tooth root geometry in order to reduce the tooth root stress should be proved. At least it should not lead to an increase in the tooth root stress. For this reason, the optimized tooth root that resulted from the calculation with the isotropic material and, exemplarily, the optimized tooth roots of materials c, d and e were calculated again with the remaining 25 material property files. Fig. 2-11 shows the proportional deviation of the tooth root stress for each material property file at the respective optimized model with respect to the original model. On average, a slight stress reduction of 1.5% was achieved for the model that was optimized with the isotropic material. However, the individual results vary between a reduction of 15% and an increase of 11.6%. The optimizations with anisotropic material property files achieved an average stress increase of 0.5, 2.4 and 8.5%. For the individual simulations, the stress reductions ranged from 16.4 to 20.5% and were larger than those for the isotropic cases, but a larger stress increase of 15–23.5% was also reached. As the average values already showed, no substantial reduction in the tooth root stress could be achieved by calculating the optimized geometries with other material property files. However, for the dimensioning of microgearwheels, the ‘worst case’ has to be assumed and here all four optimized models showed a substantial increase between 11.6 and 23.5%. It remains to note that the single material property files lead to very different stress values for different geometries. The following example demonstrates the difficulties of finding a general optimized contour for the tooth root. For two anisotropic material property files (here denoted as materials f and g), the different optimized geometries were recalculated. In the case of the optimized geometries of the isotropic material, a stress reduction of 6.6% for material f could be achieved but a stress increase of 5.2% for material g was obtained. In the case of the optimized geometries of materials c, d and e, stress deviations of –11.9, 15 and 13.1% for material f and of 14, –18.2 and 19.2% for material g were obtained, which means that these two material property files behave differently. Consequently, the evaluation of the results does not help in finding a general trend. As Fig. 2-11 demonstrates, all optimized geometries show a stress increase in at least 50% of the new calculations with other material property files. Owing to the inhomogeneous stress distribution caused by the grain structure, a general loaction point where the geometries could be varied in a way that would minimize the local stress increase cannot be identified. Although it is possible to find an optimized solution

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for each case, it was not possible to find a global solution for the different grain orientations. Therefore, the optimization of the tooth roots cannot be carried out using this method.

2.5 Conclusion

Independent of the number of grains in the cross-section that are observed in parts manufactured within the Collaborative Research Center, the limits of isotropic modeling for different materials (zirconia, 17-4 PH and gold instead of Stabilor G) were demonstrated. The microspecimens require at least eight grains in the relevant cross-section to assume quasi-isotropic material behavior. If a smaller number of grains is present in the critical cross-section, anisotropy effects must be taken into account owing to their possible stress-increasing influence. The examination of pores in microtensile specimens with few grains in the cross-section showed a deviation of the stress increase depending on the orientation of the surrounding grains. Results of further simulations have to be implemented or the influence of additional stress increases have to be considered. The demonstrator components do not deliver a limit value for the assumption of quasi-isotropic material behavior by means of the examination results. The stress deviation of the planetary gearwheel tooth simulation with seven grains in the cross-section is, according to the material, up to 40% and is therefore too large to be considered as quasi-isotropic. Here the dimensioning has to add an additional safety factor. With regard to the research activities in the Collaborative Research Center, zirconia, owing to its small grain size, can be considered as a quasi-isotropic material for the actual dimensions of the demonstrator components and microspecimens. In the case of microcomponents of 17-4 PH and gold/Stabilor G, an additional safety factor has to be included in order to account for the anisotropy of the relatively small number of grains in the cross-section of the component. The macroscopic approach of the fully stressed design was examined with regard to its application to microcomponents with few grains. Owing to the inhomogeneous stress distribution caused by the grain structure, it is not possible to find an optimized tooth root geometry without a substantial stress increase for individual grain configurations. Regarding the shape optimization, it remains to say that the geometry changes are so small that they are smaller than the tolerances of the present manufacturing processes (see Chapter 4). Therefore, today the implementation of an optimized geometry is difficult to achieve.

2.8 References

2.6 Outlook

Current calculation models concentrate on the three-dimensional modeling of grain structures in. This is the requirement for examining complex geometries and stress states, e.g. caused by pores, in grain structures. At the same time, further variation possibilities of the Poisson point process are developed by means of variation of the Hardcore value in specific component areas, in order to be able to account for changes in grain structures in different component areas (e.g. middle/edge). To reduce the extensive manual work in the model of microcomponents, solutions were found for the automated meshing of three-dimensional models. The conversion into software and the calculation of three-dimensional grain structures is in progress.

2.7 Acknowledgment

This research is financially supported by the German Research Foundation (DFG) within the Collaborative Research Center (SFB) 499, ‘Development, Production and Quality Assurance of Primary Shaped MicroComponents from Metallic and Ceramic Materials’.

2.8 References 1 N. Wellerdick-Wojtasik, Theoretische

2

3

4

5

und experimentelle Untersuchungen über die Fließflächenentwicklung bei großer Scherdeformation; Dissertation, VDI-Verlag, Vol. 18, Mechanik/Bruchmechanik 1997, 215. S. Weyer, Mikromechanisches Modell zur Ermittlung effektiver Materialeigenschaften von geschädigten Polykristallen; Dissertation, Universität Karlsruhe (TH), 2001. J. Osher, U. Lorz, Quantitative Gefügeanalyse – Theoretische Grundlagen und Anwendungen. Leipzig: Deutscher Verlag für Grundstoffindustrie, 1994. D. Stoyan, W. S. Kendall, J. Mecke, Stochastic Geometry and Its Applications, 2nd edn; Chichester: Wiley, 1995. D. Stoyan, H. Stoyan, Fraktale Formen Punktfelder; Berlin: Akademie-Verlag, 1992.

6 A. Okabe, B. Boots, K. Sugihara, Spatial

7

8 9

10

Tesselations – Concepts and Applications of Voronoi Diagrams; New York: Wiley, 1992. A. Fröhlich, Mikromechanisches Modell zur Ermittlung effektiver Materialeigenschaften von piezoelektrischen Polykristallen; Dissertation, Universität Karlsruhe (TH), 2001. H. J. Bunge, Texture Analysis in Materials Science; London: Butterworths, 1982. H. Riesch-Oppermann, VorTess generation of 2-D random Poisson–Voronoi Mosaics as framework for the micromechanical modelling of polycrystalline materials – algorithm and subroutines description; Forschungsbericht FZKA, 6325, 1999. S. Weyer, A. Fröhlich, H. Riesch-Oppermann, L. Cizelj, M. Kovac, Automatic finite element meshing of planar Voronoi

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11

12

13

14

15

16

tessellations. Eng. Fracture Mech. 2002, 69, 945–958. MacNeal Schwendler Corporation, MSC.Patran 2004, PCL Manuals; Santa Ana, CA: MacNeal Schwendler. 2004. R. Hill, Elastic properties of reinforced solids: some theoretical principles. J. Mech. Phys. Solids 1963, 11, 357–372. Z. Hashin, Theory of mechanical behavior of heterogeneous media. Appl. Mech. Rev. 1964, 17, 1–9. S. Weyer, A. Brückner-Foit, Versagensverhalten keramischer Werkstoffe unter Druckbeanspruchung. Jahresbericht 1999; Karlsruhe: Institut für Keramik im Maschinenbau, 1999. S. Weyer, A. Brückner-Foit, A. Fröhlich, Overall properties of ceramics subjected to compressive loading. In: International Conference on Engineering Ceramics and Structures; Cocoa Beach, FL: American Ceramic Society, 2000, pp. 101–107. A. Fröhlich, S. Weyer, D. Metz, O. Müller, A. Brückner-Foit, A. Albers, Investigations on the reliability of FEA calculations on the microscopic scale. In: Technical Proceedings of the 2001 International Conference on Computational Nanoscience,

17

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19

20

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Hilton Head Island, SC; 2001, pp. 161– 164. FKM Richtlinie, Rechnerischer Festigkeitsnachweis für Maschinenbauteile, 4., erweiterte Ausgabe 2002; Frankfurt/Main: VDMA Verlag, 2002. H. Gallagher, C. Zienkiewicz, Optimum Structural Design – Theory and Applications; London: Wiley, 1973. W. Karush, Minima of functions of several variables with inequalities as side conditions; MSc Thesis, Department of Mathematics, University of Chicago, 1939. H. W. Kuhn, A. W. Tucker, Nonlinear programming. In: Proceedings of the Second Berkeley Symposium on Mathematical Statistics and Probability; Berkeley, CA: University of California Press, 1951, pp. 481–492. E. Schnack, Ein Iterationsverfahren zur Optimierung von Kerboberflächen, VDIForschungsheft 589; Düsseldorf: VDI-Verlag, 1978. J. Marz, N. Burkardt, A. Albers, in: Conference Proceedings EUSPEN; 2003, Vol. 1, pp. 23–26.

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3 Modeling Micro PIM D. Kauzlaric, A. Greiner, J. G. Korvink, Institute for Microsystemtechnology, University of Freiburg, Germany M. Schulz, R. Heldele, Institute for Material Research III (IMF III), Forschungszentrum Karlsruhe, Germany Abstract Micro powder injection molding (Micro PIM) consists, in essence, of the same process steps as its macroscopic counterpart (see Chapter 10). In order to obtain a reliable and reproducible process flow suitable for mass production applications, the process parameters must be identified. It turns out that the determination of such process parameters by trial and error can be improved considerably, as has been shown for the macroscopic process, by applying computer simulations to predict the process result [1]. Commercially available simulation programs for macroscopic modeling, however, seem to fail when it comes to the prediction of the process in microscopic dimensions. The greater the complexity of the mold geometry, the less reliable the predictions with conventional simulation tools are [2]. A comparative study of different simulation tools can be found in [3]. One of the common main drawbacks of all the programs is their lack of standard rheological models for powder filled feedstocks. In fact, most of the simulation tools use single-phase models for the description of the feedstock. This does not allow for the prediction of, e.g., segregation effects. As a consequence, there is a need for new approaches in this area. For the description of the microscopic dynamics of fluid flow, molecular dynamics is one of the most popular methods. It achieves very reliable fundamental results for fluid behavior. When applied to geometries that extend above the nanometer scale, the computational costs are so high that it is not possible to simulate an injection molding process by this method in the foreseeable future. Dissipative particle dynamics (DPD) is a particle method on the mesoscopic scale. Therefore, it is a promising candidate to fill the gap between microscopic and continuum methods. This chapter explains our efforts to apply DPD to the modeling and simulation of Micro PIM and its verification by suitable laboratory experiments.

Advanced Micro and Nanosystems Vol. 3. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31208-0

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Keywords micro powder injection molding; computational fluid dynamics; dissipative particle dynamics; simulation

3.1 3.2 3.3 3.4 3.4.1 3.4.1.1

3.4.1.2 3.4.1.3 3.4.2 3.4.2.1 3.4.2.2 3.4.2.3 3.4.3 3.4.3.1 3.4.3.2 3.4.4 3.5 3.6 3.7 3.8

Introduction 52 Flow Behavior of Polymers in Microchannels 54 Dissipative Particle Dynamics 57 Matching Macroscopic Parameters with DPD 59 Matching the Equation of State 60

The Free Energy Approach for Many-body Dissipative Particle Dynamics (MDPD) 60 Warren’s Approach for MDPD 63 EOS Measurement, Modeling Pressure, Specific Volume and Temperature (pvT) 64 Fluid–Wall Interaction: the Contact Angle 67 Reflecting Boundary Conditions 68 Interactions for Reproduction of Wetting and Friction 69 Measurements of Contact Angle Between Molding Material and Mold Insert 70 Viscosity Modeling and Measurement 72 The Parameters in DPD 72 Shear Viscosity Measurements 74 Thermal Conductivity and Thermal Expansion, Measurement and Modeling 76 Test Geometries 80 Conclusion 82 Acknowledgment 82 References 82

3.1 Introduction

The modeling and simulation of the behavior of materials in various applications of modern microsystem technology is a challenging task. This is especially the case for form-giving process simulation. In practice, high costs or restricted possibilities in machine finishing require special care to be taken in the design of the mold. Simulation tools emerging from theoretical modeling that ranges from quantum physics of atomic-scale phenomena to continuum descriptions of macroscopic behavior [4] and the system behavior, are still lacking engineering tools to support the design processes in microsystem technology. In microsystems, the presence of several orders of magnitude of length scales requires the coupling of different tools appropriate for simulation at the respective spatial scale. Given these boundary conditions for the design of microsystems, the

3.1 Introduction

use of advanced computer-aided design (CAD) tools promises a reduction in the extent of physical testing necessary to prototype a device. Moreover, these looked-for tools are the auxiliary means to reduce production costs for microsystems. Activities to incorporate various physical models at different length scales already exist (e.g. [5]), which show great promise for the solution of the given task to simulate an entire process chain in microsystem technology. The simulation of complex fluidic processes in a common framework with reasonable computational effort – as is mandatory for its application in an engineering design process – might be realized by superimposing particle dynamics on a conventional computational fluid dynamics (CFD) simulation. Typically, the particle dynamic simulation will be computationally more expensive than the CFD solver. Moreover, the iterative procedure for fluid solvers, due to the non-linear nature of the Navier–Stokes equations, will force the costly particle dynamics calculation to be repeated in every iterative loop. To overcome this problem, it is desirable to improve the coupling of the microscale simulation featuring noniterative and explicit time marching to the continuum model acting on a much coarser length scale and thus eventually leading to a feasible simulation process. At the microscopic length scale, molecular dynamics (MD) represents an attractive simulation technique for CFD. However, for length scales larger than 10 nm, a coarse graining of the molecular model is desirable owing to rising computational costs. A very promising technique is DPD, first introduced by Hoogerbrugge and Koelman in 1992 [6]. In essence, it is a treatment of the dynamics of quasi-particles each representing small sets of the liquid’s molecules, by stochastic differential equations in a fashion similar to a Langevin approach. It combines features from MD and lattice gas methods (e.g. [7]). Since its first introduction, this method has been applied to the simulation of a wide range of phenomena, especially in the area of materials science. Many investigations have been carried out in order to understand the capabilities of DPD as applied to CFD problems [8–15]. DPD is a mesoscopic simulation method capable of ‘bridging the gap’ between atomistic and mesoscopic simulation [16]. Its application range for different simulation tasks includes mesoscopic dynamics of colloids [17], binary fluids and the matching of macroscopic properties with DPD [18], domain growth and phase separation in immiscible fluids [19] and the simulation of rheological properties [20]. The last is a very important characteristic of this method and augments its capability of predicting material properties, so necessary for microforming processes. Recently, DPD has experienced several improvements. Its algorithmic optimization was a major focus, thereby making it a method appropriate for application in the engineering field [21–23]. The inclusion of energy conservation in the particle–particle interaction for the set of stochastic differential equations describing a DPD model has been derived [24, 25]. This is an important extension for heat flow applications. Recently, phase change models, built on the energyconserving DPD models [26] for a solid–liquid phase transition, have been included. Liquid–vapor coexistence, instead, cannot be modeled within a standard

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DPD approach unless one drops all the advantages of the method arising from the use of larger time steps than are typical in MD [16]. Nevertheless, there are several improvements that lead to a tolerable increase in computational costs, that allow for modeling of liquid–vapor systems. The technical issues will be considered in the following sections.

3.2 Flow Behavior of Polymers in Microchannels

Accompanied by the rapid development of microfabrication techniques, MEMS technologies are now approaching a level where the standard macroscopic material properties, such as Young’s modulus, tensile strength and Poisson’s ratio, cannot be employed without restrictions at the microscopic level. This statement is also true for fluid properties, which play a large role in microinjection molding. Simulating the filling of microcavities with standard flow simulation packages such as FLOW-3D and CFX4 have shown that the calculated flow front velocities for water do not match the experimental data and differ by a factor of up to 70% [27]. This discrepancy between experimental and theoretical data can be explained by various effects that occur in microdimensions. The large surface-tovolume ratios result in strong influences from intermolecular interactions between a fluid and a surface, e.g. surface tension effects, wall slip effects, viscous effects (see Sections 3.4.2 and 3.4.3) and adhesion effects [28]. Most simulation tools designed for injection molding use the generalized Hele–Shaw flow model for non-isothermal and non-Newtonian viscous fluids [29]. This model does not take into account that, when the external characteristic length (e.g. channel depth) becomes comparable to the internal characteristic length (e.g. radius of gyration of a polymeric molecule), the long-range intermolecular forces between the polymer chains and the molding surface begin to influence rheological and capillary behavior significantly. Various experimental observations have shown that the rheological behavior of fluids near rigid surfaces is significantly altered compared with the bulk behavior measured at a sufficient distance from surface walls [30–34]. It has been shown, for example, that the viscosity near a channel wall is 50–80% higher than in the bulk fluid. The surface–fluid interactions are long-range intermolecular forces that cause orientational effects. Based on that, Eringen and Okada [33] developed a nonlocal continuum theory of viscous fluids with the following viscosity model: " g ˆ gb 1 ‡ c

 2 # Rg D

…1†

Equation (1) describes the fluid viscosity g, where gb is the bulk viscosity, c is a dimensionless constant, Rg is the radius of gyration for fluid molecules and D

3.2 Flow Behavior of Polymers in Microchannels

is the external characteristic length, e.g. the channel diameter. The gyration radius Rg is defined as the root mean square of unperturbed random recoil of polymeric molecules and depends strongly on the molecular weight of the polymer [35]. If D and the radius of gyration are of the same order of magnitude, the viscosity will rise significantly. Fig. 3-1 shows this relationship for a Newtonian fluid based on the coefficients for polystyrene–cyclohexane [32], with c = 19.5 and Rg = 26 nm. A significant viscosity change near the surface wall can be observed. Possibly this effect influences the filling of structural details during micro powder injection molding. The rheological data measured in macroscopic experiments and used in standard simulation packages for injection molding are not completely suitable for modeling the flow in microchannels. The most common simulation tools for injection molding, MoldFlow and CMold, both assume no-slip boundary conditions at the mold walls [36]. Various theoretical studies [37–40] have attempted to describe the wall-slip phenomenon with some success, although most of them are not properly derived from fundamental physical principles. Yao and Kim [41] stated two possibilities as to why slip can more easily occur in microchannels. The linkage of polymer chains with each other can, on a basic mechanical level, be described with a spring– damper system. If the channel size decreases, the number of springs across the gap also decreases, and therefore the polymer melt appears to be more rigid, resulting in wall slip. They also showed that an increased amount of pressure drop is necessary to fill microchannels. The consequence is a higher pressure on the fluid near the wall, which can lead to either a separation of grafting polymer chains and the surface or disentanglement of grafting chains with bulk chains, thus resulting in wall slip. In practice, this can lead to a plug flow in microcavities with a nearly homogeneous flow velocity profile.

Fig. 3-1 Rise of viscosity near solid surfaces predicted by Eringen and Okada [33]

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This observation leads to another important fact for channels in microdimensions. A significant effect when scaling a fluidic system into microdimensions can be seen by looking at the Reynolds number: Re ˆ

vD m

…2†

where v is the fluid velocity, D the capillary diameter and m the kinematic viscosity. The Reynolds number relates the inertial forces to the frictional (viscous) forces and will tend towards low numbers if the viscous forces dominate. It describes the transition point between laminar (low Reynolds numbers) and turbulent (high Reynolds numbers) flow. With small cavities in the micrometer range and polymer viscosities high above 1 Pa s, it is safe to assume laminar flow during Micro PIM for simple geometries (Re <<1). The influence of surface tension and the resulting contact angle between molding surface, ambient atmosphere and feedstock is yet another fact that today’s simulating tools ignore. Some microdevices are based on the fact that with decreased channel size, an appropriate fluid can fill a whole capillary system without external pressure [42]. This is attributed to the capillary pressure as described by the Washburn equation: Dp ˆ

4S cos W d

…3†

where Dp is the equivalent pressure change at the front of a capillary flow, due to the surface tension S, W is the contact angle of the fluid with the capillary wall and d is the diameter of a capillary with circular cross-section. With the experimental data in Section 3.4.2, the influence of capillary pressure on mold filling during Micro PIM can be estimated. The contact angle of

Fig. 3-2 Calculated capillary pressure of a binder system on molding surface with a contact angle of 368 and surface tension of 23 mN m–1 at 160 8C

3.3 Dissipative Particle Dynamics

the binder-system on the molding surface at 160 8C is about W = 368 and the corresponding surface tension S = 23 mN m–1. The resulting pressure is illustrated in Fig. 3-2. Typical filling pressures of several hundred MPa are used in l PIM so that the influence of wetting seems to be negligible [41]. In addition, it is necessary to determine the compression ratio in the microcavities since the expected pressure loss is likely to be high.

3.3 Dissipative Particle Dynamics

Continuum-based computational fluid dynamics (c-CFD) uses the available discretization methods, such as Finite Elements, Control Volume or Finite Differences, to treat numerically the partial differential equations of motion that describe the hydrodynamic behavior of a fluid. The best known among these are the Navier–Stokes equations. In the DPD method, one discretizes the continuum fluid into small portions as shown in Fig. 3-3. Usually, these portions are represented by point particles. One might think of a DPD particle as a representation in terms of the center of mass dynamics of a fluid portion. The positions and momenta of the particles are updated in continuous phase space at discrete time steps. The updates are computed by applying Newton’s second law for a particle of mass mi: X dr i p dp ˆ i ; i ˆ F ext f ij i ‡ dt mi dt j6ˆi

…4†

where ri and pi are the position and momentum vectors of particle i and F ext is i R C an external force field acting on each particle. f ij ˆ F D ‡ F ‡ F is a net pair ij ij ij R force acting between two particles i and j, where F D ij is a dissipative force, F ij is C a stochastic force and F ij is a conservative force. R Suitable forms for F D ij and F ij , which fulfil the necessary condition of Galilean invariance, are FD ij ˆ

cwD …r ij †…eij
vij †eij

…5†

Fig. 3-3 Schematic picture of a DPD discretization of the continuum (from left to right) or a coarsening of the atomistic picture (from right to left)

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and F Rij ˆ rnij wR …r ij †eij

…6†

The vector eij = r ij /r ij is the unit vector pointing from particle j to particle i. vij ˆ vi vj is the relative velocity between the particles. nij is a random number with the properties hnij …t†i ˆ 0 and hnij …t†nkl …t0 †i ˆ …dik djl ‡ dil djk †d…t t0 †, where Rdij ˆ 1 for i = j and dij ˆ 0 for i 6ˆ j, and d…x† is the Dirac d-function with d…x†dx ˆ 1. The scalar c can be interpreted as a friction coefficient and r is a scalar noise amplitude. wD …r ij † and wR …r ij † are weight functions that determine the range of the forces and their strengths as a function of the interparticle distance r ij . For r ij  rc, where rc is the cutoff distance, these functions vanish. For a well-defined equilibrium temperature, detailed balance requires a fluctuation-dissipation theorem for DPD to be fulfilled [43], which is wD …r ij † ˆ wR2 …r ij †

…7†

and 2kB T c ˆ r2 m

…8†

where m is the mass of a particle, kB is the Boltzmann constant and T is the equilibrium temperature. This means that the stochastic and dissipative forces act together as a momentum-conserving thermostat. Usually, the weight functions are chosen to be ( wD …rij † ˆ

wR2 …rij †

ˆ

1 0;

 r ij 2 ; rc

r ij < rc rij  r c

…9†

A convenient choice for the conservative force is [16] F Cij ˆ awC …rij †eij

…10†

with a > 0 representing a repulsion amplitude and wC …rij † ˆ wR …rij †. For the numerical integration of Equation (4), the modified velocity Verlet algorithm is used, as described in [16]. It takes into account that the dissipative force depends on the velocities and vice versa. When using the alternative Lowe–Andersen thermostat (see Section 3.4.3) later on, the standard velocity Verlet algorithm known from MD is sufficient, since the thermalization is no longer coupled to the force computation. Because of the non-systematic coarse-graining procedure of DPD, quantities such as length, time or energy lose their direct link to the corresponding quantities of real fluids. For the DPD algorithm, this poses no fundamental problem, since, as for MD, one can always compute using reduced units. This is what

3.4 Matching Macroscopic Parameters with DPD

was done for all the DPD-based simulation results presented in this chapter. The main difference from MD is that the direct conversion from reduced to real units is lost. This has to be re-established by some additional empirical assumptions, such as an agreement on how many MD atoms make up one DPD particle. Additionally, the connection to real units can be re-established by matching important properties of real fluids, such as the equation of state (EOS) or the viscosity. How the matching was performed will be shown in more detail in the following sections.

3.4 Matching Macroscopic Parameters with DPD

Both c-CFD and DPD need experimental thermodynamic and rheological data. The most important information includes the EOS of the feedstock system, the viscosity including its dependence on temperature and shear rate and the thermal conductivity and heat capacity for the feedstock and the mold material. Additionally, one should determine whether surface effects are relevant. This includes measurements of the surface tension and of contact angles for feedstock on mold material and binder on powder material. There is a fundamental difference between c-CFD and DPD in the way in which experimental data are used. For c-CFD, experimental datasets are input, and they are used to set up material parameters for the underlying partial differential equations through constitutive equations or look-up tables. In DPD, properties of a fluid, such as the viscosity, are not specified directly, but emerge from the underlying particle interactions of the model. In this case, the experimental measurements are used to tune the interparticle interactions, until experiment and simulation match. Table 3-1 shows examples of phenomena that are well suited for modeling with DPD. In order to reproduce these, detailed tuning of the parameters that

Table 3-1 Application fields of DPD and actions needed in order to reproduce the phenomenon

Effect

Challenge

Solution

Flow

Reproduce flow with particles

Phase boundary

Reproduce a flow front

Wetting and nonwetting Heat conduction

Surface and interface energies Transient thermal behavior including boundary Reproduce multi-phase flow with particles

DPD, interparticle forces as in Equations (4), (5), (9) and (10) MDPD, interparticle forces depend on density Fluid–wall interaction forces Energy-conserving DPD

Mixing and segregation

Introduce discriminative particles

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enter DPD must be performed. The following sections describe this matching process for specific fluid properties. Section 3.4.1 deals with the reproduction of the EOS. A discussion of static surface effects from an experimental and modeling point of view follows in Section 3.4.2. Section 3.4.3 reports on experiments and simulation for viscosity matching and Section 3.4.4 presents thermal experimental results and gives a short overview of possibilities for reproducing the correct thermal behavior in DPD. 3.4.1 Matching the Equation of State 3.4.1.1 The Free Energy Approach for Many-body Dissipative Particle Dynamics (MDPD) In DPD, the EOS can be ‘measured’, which means it can be computed. This is done by fixing the temperature and density in a periodic system of particles and by measuring the pressure. Several techniques exist for measuring the pressure [16]. The virial expression for the pressure in three dimensions gives

1 p ˆ qkB T ‡ 3V

*

X

+ r ij
f ij

…11†

i<j

where V is the volume of the simulated system and the sum indicates a sum over all pairs of particles. The average is a time average over the time steps after having reached equilibrium. For a given choice of stochastic and dissipative forces producing a well-defined equilibrium temperature T, the standard DPD model, as described above, can be found to collapse to the following EOS [16]: p ˆ qkB T ‡ aaq2

…12†

where a > 0 is a constant and a > 0 is the repulsion coefficient of the conservative force defined in Equation (10). This shows that, for adjusting the EOS, the conservative force is the only free parameter. Now, the key question is whether this EOS fulfils our needs concerning the simulation of Micro PIM. One requirement, which is definitely mandatory, is the correct formation of a flow front, which means that we require the emergence of an interface between the liquid and a sort of gas phase. From Equation (12), it can immediately be seen that, for any temperature T, a single phase with one equilibrium density will exist. Fig. 3-4 shows the consequences for a flow simulation over a backward-facing-step geometry. The walls were modelled by using Maxwellian reflection. This will be described in Section 3.4.2. The snapshots show that, if liquid–vapor coexistence is not possible, no stable flow front can be established and the DPD fluid behaves more like a gas. Particles on a free surface spread away from it.

3.4 Matching Macroscopic Parameters with DPD

Fig. 3-4 Flow of standard DPD fluid through a backward-facing step

It is obvious what has to be done in order to obtain liquid–vapor coexistence. In MD, the simple Lennard–Jones potential can produce liquid–vapor coexistence on an atomistic scale. One of the consequences of the coarse graining procedure in DPD is the removal of the attractive part of the conservative force. It should be clear that, at a free surface, such a potential leading to an interparticle conservative force as given in Equation (10) is not able to attract particles to the bulk of the fluid in a way that the surface remains stable. Actually, it is possible to define purely repulsive interaction potentials which can produce liquid–vapor coexistence by making the interaction range dependent on density [44]. The intuitive solution is, of course, to add an attractive force. Pagonabarraga and Frenkel [45] have developed a general procedure, which starts out from the desired EOS that should be reproduced. For this purpose, the free energy F of the system is expressed as a sum of the individual free energies w…qi † of the DPD particles: Fˆ

X

w…qi †

…13†

i

where qi is a local density approximation for particle i. The free energy is the sum of a kinetic and an excess part. The variation of the excess part wex can be taken to define the force acting on a particle by Fi ˆ

@ X ex w …qj † @r i j

…14†

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The local density can be defined as qi ˆ

X

w…r ij †

…15†

j

where w…r ij † can be chosen in a similar way as the weighting functions for the forces are defined [see Equation (9)]. From Equation (4) and since random and dissipative forces may not contribute to the pressure, the force is F i ˆ F Ci ˆ

X

…16†

F Cij

j

which, together with Equations (14) and (15), leads to "

F Cij

# @wex …qi † @wex …qj † @w…rij † ˆ ‡ eij @qj @qi @rij

…17†

for the conservative pair interaction. Now, the correct conservative interaction for our system can be determined by matching the desired equation of state to the thermodynamic definition of the pressure: p ˆ qkB T ‡ q2

@w @q

…18†

and obtaining an expression for @w/@q. For example, for the EOS pvdW ˆ

qkB T 1 bq

…19†

aq2

of a van der Waals fluid, one finds @w bkB T ˆ @q 1 bq

…20†

a

which leads to " F Cij

ˆ

bkB T 1 bqi

 a ‡

bkB T 1 bqj

!# a

@w …rij †eij @rij

…21†

One can see that, generally, the resulting forces are density-dependent manybody forces. Trofimov et al. [46] determined that the computed EOS deviates slightly from the desired one. An improvement was achieved by iteratively determining a ‘better’ local density approximation, which minimizes the difference between the

3.4 Matching Macroscopic Parameters with DPD

average hqi i and the global density n=N/V, where N is the total number of DPD particles in the system and V the total volume.

3.4.1.2 Warren’s Approach for MDPD Warren [47] directly modified the repulsion coefficient a of the original conservative force from Equation (10), which consequently also changed the EOS. The simple quadratic dependence on density was modified by making the repulsion coefficient density dependent. More precisely, the force from Equation (10) was made attractive by choosing a < 0 and adding a repulsive many-body force of the form

F Cij ˆ B…qi ‡ qj †wC …rij †eij ; B > 0

…22†

with a different cutoff range rd < rc and w…rij † ˆ

 15 1 2prd3

rij rd

2 …23†

for the computation of the local densities of a 3D DPD fluid according to Equation (15). For this computation, the same cutoff distance rd was used.

Results In [47], the above was tested successfully for the static case of a pending droplet. Here, the dynamic test case is investigated. One of the parameter sets from [47] was taken (a = –40, B = 40, rc = 1, rd = 0.75) and, for temperature adjustment, the Lowe–Andersen thermostat [22] was used, instead of the dissipative and stochastic forces described in Equations (5)–(7). This thermostat will be described in detail in Section 3.4.3. The step geometry of Fig. 3-4 was considered again, in order to test whether stable flow fronts can be obtained in a dynamic non-equilibrium and non-stationary situation. Fig. 3-5 shows that this feature was successfully achieved. Two conclusions can already be drawn at this stage. First, since the flow front emerges naturally from inter-particle interactions, DPD is a more natural flow solver for fluids with free surfaces than standard c-CFD, where the flow front has to be tracked in order to determine, for example, which viscosity has to be assigned to a certain flow region. Second, already this simple test case allows for a qualitative prediction of the jetting effect, i.e. the free surface flow of the mold mass without contact with the mold’s wall. In the example, this feature is clearly visible. It turns out that halving the driving force reduces the jetting effect considerably. The effect never disappears completely. Whether it leads to severe failures in the molded micropart depends on the length and geometry of the remaining flow path. Reducing

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Fig. 3-5 Flow of MDPD fluid over a backward-facing step

the driving force to one-tenth of the original value leads to an arrest of the fluid flow at a position only a few multiples of rc behind the inlet, i.e. in the narrow part of the geometry. This means that short shots can already be observed. The next step is to take the experimentally determined Tait parameters, to find a polynomial approximation and to determine the corresponding conservative DPD interactions from it following the procedure described above. Then, the DPD-EOS has to be computed in order to test how well the original EOS has been reproduced. If this test is successful, the DPD units can be synchronized with the real units, based on this thermodynamic matching. A similar synchronization can be performed based on a rheological matching, which means on a matching of the viscous flow behavior (see Section 3.4.3). It was pointed out in [47] that the gas phase is very dilute. This is reasonable, since it represents the pure gas phase of the liquid material, which one would only approximately obtain by equilibrating a droplet in high vacuum. In l PIM the material is also injected into vacuum, but not high vacuum. This means that one has to check if additional DPD particles of a second species representing the air have to be added to the simulation. However, their presence is definitely less important than for macroscopic injection molding, where one typically does not evacuate the mold.

3.4.1.3 EOS Measurement, Modeling Pressure, Specific Volume and Temperature (pvT) The interrelation between the state of matter data, i.e. pressure p, specific volume v and temperature T, allows for the prediction of a material’s behavior under injection molding conditions. Different methods exist for the determination of the data for the isobaric pvT plot, where the specific volume v, i.e. the reciprocal density, is plotted against temperature at constant pressure values. Material properties such as compressibility and phase transformations are responsible for non-linear characteristics of the diagrams. Depending on the method used for determination, the pvT data can also serve for thermal expansion calculations

3.4 Matching Macroscopic Parameters with DPD

of the material during the filling phase or volumetric shrinkage calculations during the packing phase of the injection molding cycle. The state of matter diagram can be measured, for instance, with a high-pressure capillary rheometer. A special nozzle is used to realize a closed system filled with material. The pvT test is performed on cooling from the processing temperature to well below the ejection temperature to avoid any decomposition at this time. Typically the measurement starts from the molten state at normal processing temperature since it is necessary to make sure that no voids are enclosed in the molding material. Data are logged on a series of isothermal runs while the pressure is raised to several hundred MPa for each isotherm. The measurement of the specific volume is subsequently repeated during heating to the maximum processing temperature, since decomposition at this time will not further influence the data. A schematic plot of specific volume versus temperature for both a crystalline and an amorphous material is shown in Fig. 3-6. The pvT relationship is typically modeled using a modified two-domain Tait equation [54–56]:  V…T; p† ˆ V0 …T† 1

 C ln 1 ‡

p B…T†

 …24†

where V(T, p) is the specific volume at temperature T and pressure p, V0 is the specific volume on the zero gage pressure isobar, C = 0.0894 is a universal constant for this model [57], B is the pressure sensitivity of the material. Two temperature domains, one greater than the transition temperature Ttrans and one less than Ttrans , are required to model the relationship because the thermodynamic properties of polymers change at this temperature. Ttrans at zero gage pressure is denoted b5. The specific volume obtained by extrapolating the zero-pressure isobar curve to the transition temperature is denoted b1. This value is the same for both domains when crossing the glass transition. When the material is semicrystalline, however,

Fig. 3-6 Typical transitions in pvT measurement, scheme of determination of the Tait parameters for (a) amorphous and (b) crystalline transitions

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the transition due to crystallization is accompanied by an abrupt change in specific volume, such that b1m, the melt specific volume at Ttrans on the zero-pressure isobar, is greater than b1s. For each domain, the rate of change of specific volume with temperature at zero pressure is measured by the respective b2 constants: V0 ˆ b1m ‡ b2m …T V0 ˆ b1s ‡ b2s …T

b5 † for T > Ttrans b5 † for T < Ttrans

…25† …26†

The remaining constants, b3 and b4, characterize the pressure sensitivity B of the material: B…T† ˆ b3m exp‰ b4 …T

b5 †Š

for T > Ttrans

…27†

B…T† ˆ b3s exp‰ b4s …T

b5 †Š

for T < Ttrans

…28†

In particular, a positive b4 reflects a specific volume which becomes more pressure sensitive with increasing temperature. Typically, the binder for l-PIM consists of more than one thermoplastic component (see Chapter 11.2). Therefore, the pvT plot shows more than one transition temperature. For the modeling and simulation of the molding step, the typically observed domains are the solid state (at temperatures below the first Ttrans;1 ) and the state where all binder components are completely molten (above Ttrans;2 ). To validate a two-domain Tait equation for complex feedstock systems in a c-CFD simulation tool, it is necessary to skip the temperature range between the first and the last Ttrans where the material is only partly molten. The value for b5 is then defined by the intersection of the curves based on the values beyond the first Ttrans;1 and above the highest Ttrans;2 (see Fig. 3-7). Because of

Fig. 3-7 Determination of the Tait parameter b5 for multiphase compounds, idealized v,T diagram

3.4 Matching Macroscopic Parameters with DPD

this approximation, the cooling effects during the injection phase cannot be described accurately by using the mentioned two-domain Tait model on multiphase compounds. Therefore, considerations have to be made on extending the model, applying it repeatedly or by developing a multi-phase model. In [57] and [58] an extended non-equilibrium Tait model is described, introducing further terms Vt(p,T) with additional parameters. 3.4.2 Fluid–Wall Interaction: the Contact Angle

In addition to fluid–fluid interactions, it is important to reproduce correctly the fluid interaction with the mold wall. Computationally, three types of fluid–wall interactions can be distinguished. First, there is simple mechanical interaction, which means that the fluid is prevented from penetrating a certain region delimited by a solid wall. Second, there is friction between the liquid and the wall. This usually leads to the so-called no-slip condition at the wall, which means that, directly at the wall, the parallel velocity components vanish. At high shear rates, this no-slip condition is often violated. Third, in the microfluidic regime, capillary effects can become relevant. It still has to be determined whether this is also the case for l-PIM. Estimates for l-PIM indicate that, for microstructures with diameters below 100 lm, capillary effects become important [48]. These microscale surface effects can either hinder or boost the flow, depending on the contact angle between the feedstock and the mold material. On the other hand, in [41] it is argued that surface tension, the driving force for the capillary effect, can be neglected for polystyrene, with channel sizes over 1 lm and for typical pressures (several MPa to 200 MPa) applied in polymer injection molding. The basis of their reasoning is the Washburn Equation (3). For the observed fluid (polystyrene, S=0.033 N m–1), Dp is negligible compared with the applied pressures during injection molding, for channel sizes down to less than 1 lm. As mentioned in Section 3.2, the possibility cannot be excluded that capillary effects play a role in l-PIM because of high pressure loss. In the following, some possibilities for modeling these effects will be shown. Similarly to the experimental domain, one can distinguish effects of fluid– wall interaction in the modeling domain. In c-CFD, one usually defines a slip or no-slip condition for hydrodynamics and a surface energy between liquid and vapor and a contact angle between liquid and solid for hydrostatics. In DPD, all fluid–wall models are dynamic in nature. Static properties are obtained as solutions of equilibrated particle dynamics simulations. One can define different kinds of reflection mechanisms for particles hitting a wall and, additionally, interactions can be introduced between DPD particles and the wall. Now, the chosen model and some possible alternatives for modeling the reflection mechanism are shown, and then the introduction of additional interactions is discussed.

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3.4.2.1 Reflecting Boundary Conditions For our simulations, including those described in Section 3.4.1, Maxwellian reflection [49] was applied. This means that a particle hitting a wall is reflected back into the system with a new velocity drawn from a Maxwell distribution. This distribution corresponds to a temperature which is pre-defined for the wall. Therefore, the wall acts as an additional thermostat. For a velocity vector v in 3D, the new velocity can be obtained by drawing the perpendicular velocity component vp from a Rayleigh distribution:

!

mv2p

m y…vp † ˆ vp exp kB T w

…29†

2kB Tw

and both tangential components vt from a Gauss distribution centered at zero: y…vt † ˆ …2pmkB Tw †

1 2

 exp

mv2t 2kB Tw

 …30†

This boundary condition was used without any additional interactions. The effect can be seen in Fig. 3-8. Here, results are presented for a simulation where a gravitational force was added to the system, and which points perpendicularly to the wall acting as a surface. The wall temperature and the fluid temperature were both set to T = 1. The droplet can be obtained, for instance, by starting with a cuboid or cylindrical particle distribution which does not fill the whole simulation domain. In our case, the droplet contains around 500 DPD particles. A few additional DPD particles form the dilute gas phase. Very strong de-wetting behavior can be observed, which is reasonable because of the purely repulsive interaction. Additionally, it is interesting to observe that, even for a wall temperature T ? 0, the contact angle is virtually unaltered.

Fig. 3-8 Around 500 DPD particles forming a 3D droplet with large contact angle. A gravitational force and a wall at the bottom were included. The wall was modeled by Maxwellian reflection only. Additionally, a vapor phase can be observed

3.4 Matching Macroscopic Parameters with DPD

Alternative, non-thermalizing reflection mechanisms are specular or bounceback reflection [50]. Neither changes the de-wetting behavior because they still produce purely repulsive fluid–wall interactions.

3.4.2.2 Interactions for Reproduction of Wetting and Friction In order to obtain contact angles smaller than 908, it is necessary to introduce attractive interactions between the fluid particles and the wall. A straightforward method is the introduction of frozen particles at the wall [6]. The term ‘frozen’ denotes that these particles remain fixed at their initial positions, which means that forces acting on them are ignored. However, the wall particles may exert a force on the fluid particles. Wetting behavior should be obtained if, for instance, the same MDPD interactions are applied to the fluid–wall interaction, which were already used for the fluid–fluid interaction in Section 3.4.1. An example is shown in Fig. 3-9. Three stages of a spreading droplet on a surface can be seen. The gravitational force is the same as before, and the wall has been modeled by applying Maxwellian reflection and frozen wall particles with a density of q = 6. The frozen wall particles, which are situated below the line, are not shown. For this setup, even complete wetting has been achieved. This enables us now to fix a desired contact angle between the two extremes of Figs. 3-8 and 3-9 by adjusting the interaction forces between fluid and wall particles. The thin film of DPD particles at the edge of the droplet, which can be observed during wetting, is reminiscent of the well-known precursor film of real spreading droplets (e.g. [51]).

Fig. 3-9 Around 500 DPD particles forming a 3D droplet which shows complete wetting. As in Fig. 3-8, a gravitational force and a wall at the bottom were included in the simulation. The wall was modeled by Maxwellian reflection and frozen wall particles with density q = 6. For the fluid–wall interactions, the same MDPD forces as for the fluid–fluid interactions were applied

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Alternatively to frozen wall particles, it is possible to obtain integral expressions for the force exerted by the wall on a particle at a certain position. For infinitely extended flat walls and standard DPD forces (see Section 3.3), these expressions were derived in [50]. For arbitrary wall geometries, it is probably more convenient to compute the force field from frozen wall particles, which can be removed afterwards. This saves some computational time since the forces are computed only once and stored in a look-up table for small cubic boxes in the interaction range of the walls. For MDPD forces, this procedure is less straightforward, since the force from the wall acting on the fluid particle also depends on the local density of the fluid particle itself, which is not known beforehand. If the local density of wall particles should also include contributions from the fluid particles, this is also only known at runtime. The latter could be circumvented by defining an average density of the wall, which hence is independent of the state of the fluid. Whereas slip or no-slip conditions are explicitly defined in c-CFD, they emerge indirectly from the fluid–wall interactions in DPD. Currently, the best way to obtain no-slip conditions in DPD is the combination of specular reflection with either integral or discrete fluid–wall interactions. The question remains of whether it is really desirable to impose no-slip on our system since, in this way, it is impossible to determine any transition from no-slip to slip at a certain shear rate. Usually, one assumes no-slip in l-PIM modeling. However, since high shear stresses and velocities occur, this assumption is highly questionable.

3.4.2.3 Measurements of Contact Angle Between Molding Material and Mold Insert For the characterization of contact angles, most measurements in practice are performed with a sessile drop on a planar surface. In this case, the contact angle of the polymeric binder on the solid surface is a measure for the intermolecular interaction between these two materials. The methods used to determine contact angles are not only crucial for the fabrication of compound materials. During an injection cycle, the capillary pressure influences the filling of small cavities. According to Equation (3), the pressure difference Dp in a thin capillary depends on the surface tension S of the fluid, the contact angle W between fluid and capillary surface and the diameter d of the capillary. With decreasing component dimensions, and thereby smaller structures in a molding insert, the capillary pressure rises and lessens the necessary filling pressure. Both effects are parameters which can be implemented in a simulation based on DPD. Young derived an equation for sessile drops which specifies the equilibrium of forces of surface and interfacial tensions cSV , cLV and cSL that are linked together by the contact angle:

cLV cos W ˆ cSV

cSL

…31†

3.4 Matching Macroscopic Parameters with DPD Fig. 3-10 Contact angle at sessile drop in force equilibrium

The contact angle can be measured in the three-phase region (see Fig. 3-10) consisting of the solid S, the applied liquid L and the ambient gas phase V. One of the most widely used measuring methods for contact angles is based on an optical contour analyzing system. The fluid drop rests on an exchangeable planar surface in an optional furnace which controls the ambient temperature and atmosphere. The contact angle can be measured manually with a goniometer eyepiece or digitally with a video system combined with suitable software. The final binder-powder compound that is used as the l-PIM feedstock cannot be measured with this method because of its high viscosity and inhomogeneous melting behavior. However, the contact angle of a feedstock on a surface is primarily defined by its fluent component and therefore the contact angle of the binder should match that of the compound to a good approximation. Fig. 3-11 shows, as an example, the contact angle of two different binder components on a steel surface matching the mold insert material. However, the inhomogeneous melt behavior of binder B causes large deviations. It is obvious that both binder components spread on the solid material. This characteristic, combined with a decreased binder viscosity at elevated temperatures, permits improved feedstock fabrication.

Fig. 3-11 Contact angle of two different binder components on a steel surface. Measured using a G10 Contact Angle Meter (DSA10) from Kruss AG

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3.4.3 Viscosity Modeling and Measurement

Both experiment and DPD need a measurement setup to determine the transport parameters, e.g. the viscosity. In DPD a viscosimeter is clearly an algorithm that has to be added to the program code. From MD, these kind of measurements are well known. By adjusting the parameters in DPD, the results from experimental measurement have to be reproduced. Note that the adjustment of the DPD parameters may not allow for the matching of all macroscopic variables at once.

3.4.3.1 The Parameters in DPD Once the EOS is adjusted by the conservative interactions, the rheological behavior, which means the viscosity, can be fixed. Viscosity can be adjusted by modifying the dissipative force. This can be seen from the transport properties, which were derived for the limit Dt ? 0 in [53] by solving a Fokker–Planck– Boltzmann equation for DPD using the Chapman–Enskog method. Since the stochastic force depends on the dissipative force according to Equation (7), it is the thermostat that influences viscosity. This also means that, if the range of viscosities that can be reached with the current thermostat is unsatisfactory, another thermostat may be used without distorting the previously obtained EOS. In fact, one disadvantage of the thermostat, consisting of the dissipative and stochastic forces described above, is the low Schmidt number (Sc) that it produces. The Schmidt number is the ratio of the kinematic viscosity to the diffusion coefficient, which is of the order of 1 for this thermostat. However, for liquids, Sc should usually be larger, for instance around 1000 for water. Therefore, it would be convenient if it were possible to increase the viscosity. This is also useful if the DPD fluid should represent a relatively viscous PIM feedstock. Additionally, this would increase the time-scale of the system. Another disadvantage is the non-trivial integration of the emanating equations of motion as already mentioned in Section 3.3. Both drawbacks can be overcome by applying the Lowe–Andersen thermostat [22]. In this algorithm, it is decided with a probability of CDt whether to thermalize the relative velocity vij of a given pair of particles with a particle to particle separation rij < rc . The value C plays a similar role to the friction coefficient c in Equation (5). By construction of the algorithm, C has an upper limit of 1/Dt or, conversely, there is a minimum time of 1/C which is needed for complete thermalization of the system. Therefore, larger friction coefficients not make sense. However, in practice, the same limitation exists for the dissipative and random forces of the original thermostat, since there exists an upper limit of the absolute value of a force acting on a particle due to the limited accuracy of any integration algorithm. The new relative velocity v 0ij is drawn from a Maxwell distribution with varip ance fij 2kB T=m. Then, the new particle velocities become

3.4 Matching Macroscopic Parameters with DPD

v 0i ˆ vi ‡ Dij ; v 0j ˆ vj

Dij

…32†

where 1 Dij ˆ eij …v 0ij 2

vij †
eij

…33†

Results A rough comparison was made of the accessible viscosity ranges for the three cases of standard DPD, MDPD with the original thermostat and MDPD with the Lowe–Andersen thermostat. Each simulation was performed in a periodic box with 1728 randomly arranged particles at a density of q = 5 and a temperature of T = 1. The time step was set to Dt = 0.02. Note that, as already mentioned in Section 3.3, q = 5, T = 1 and Dt = 0.02 are all given in reduced dimensionless units. The viscosity was computed via the transverse current correlation function

C…t; k† ˆ

1 hj…0; k†j…t; k†i hj2 i

…34†

where j…t; k† is a shear wave oscillating, for example, in the y-direction along the x-axis: j…t; k† ˆ

X

y

vi …t† exp‰ikx…t†Š

…35†

i

The kinematic viscosity can be obtained from the decay of the transverse current correlation function [53]: C…t; k† ˆ exp…k2 mt†

…36†

For a measurable decay, the smallest k fitting in the periodic box was imposed with a suitable amplitude as an initial condition on the system. Standard DPD with the repulsive conservative force from Equation (10) and a = 40 reaches m = 0.31 for r = 3 (c = 4.5). If the conservative force is replaced by the MDPD interactions also used in Section 3.4.1, the viscosity increases to m = 1.31 for the same c. Reducing r to 0.5 (c = 0.125) results in m = 1.14, but with a temperature deviation of roughly 20% from the desired value. This means that the friction coefficient is too low for the thermostat to work properly. Increasing r to 9 (c = 40.5) gives m = 2.50. With the chosen time step, a further increase produces numerical instabilities in the integration algorithm. Therefore, the latter value can be seen as an upper limit for this thermostat at Dt = 0.02. For the Lowe–Andersen thermostat, the probability CDt was varied between 0.0625 and 1. The former means that, on average, 6.25% of the particle pairs with rij < rc are thermalized, whereas in the latter case, every pair in the interac-

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3 Modeling Micro PIM 25

kinematic viscosity [reduced units]

74

20

15

10

5

0

0

0.2

0.4

0.6

0.8

1

dissipation probability Fig. 3-12 Kinematic viscosity of the MDPD fluid with Lowe– Andersen thermostat plotted against the dissipation probability CDt. All data points are averages over only 10 simulations. This explains the large error especially for high dissipation probabilities

tion range is thermalized. Since the time step was kept constant at Dt = 0.02, the probability range corresponds to a range of C = 3.125–50. For all simulations, the thermostat was able to keep the temperature constant at the desired value T = 1. The results are plotted in Fig. 3-12. It can be seen that the Lowe–Andersen thermostat achieves viscosities between m = 3.02 and 17.27, which is indeed larger than what is obtainable with the random and dissipative force.

3.4.3.2 Shear Viscosity Measurements The viscosity was measured using a high-pressure capillary rheometer, as e.g. sketched in Fig. 3-13. The piston of the rheometer presses the material at a constant temperature and flow rate Q through a cylindrical die of known length l and diameter d. The apparent shear rate on the wall of the tube is defined by the equation

c_ a ˆ

32Q pd3

…37†

The wall shear stress sw can be calculated from the measured pressure loss Dp across the die [62]: sw ˆ

Dp l 4
d

…38†

Therefore, the apparent shear viscosity ga can be calculated from the measured values of Q and Dp and from the geometric data for the die:

3.4 Matching Macroscopic Parameters with DPD

Fig. 3-13 Cross-section of a high-pressure capillary rheometer: 1, test chamber; 2, test flow cylinder; 3, mass pressure probe; 4, full circle capillary; 5, capillary uptake; 6, input for temperature probe; 7, capillary nut; 8, cap closure; 9, Pt100 temperature sensor; 10, hexagonal screw; 11, band heater zone 3; 12, band heater zone 2; 13, band heater zone 1; 14, area heater zone 1; 15, pressure hole

ga ˆ

sw c_ a

…39†

The data from the viscosity measurements are fitted to the Cross model [63], which handles both the Newtonian and the power law flow regimes found in polymer rheology. The model treats viscosity as a function of temperature, shear rate and pressure [Equation (40)]. Compared with other models in current use, the constants of the Cross model have physical significance. The transition between two regimes is characterized by the shear stress level at which shear thinning starts. The slope of the power law curve is characterized by (1–n). The parameter ~s corresponds to the shear modulus and g0 is the viscosity in the region where the Newtonian behavior predominates.

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However, as the bulk temperature approaches the solidification temperature, this model underestimates the viscosity rise with decreasing temperature. Therefore, the corresponding measurement standard ASTM D 3835 [64] requires the Weissenberg–Rabinowitsch correction [60], which considers the difference between apparent and actual shear rates, and the Bagley correction for the entrance pressure loss: g…T; c_ a ; p† ˆ

g0 …T; p†   g c_ 1 1‡ 0 a ~s

n

…40†

3.4.4 Thermal Conductivity and Thermal Expansion, Measurement and Modeling

Thermal conductivity is a quantity that describes the ability of a homogeneous material to transfer heat by conduction. This quantity is a transport parameter that enters a distributed parameter description of a material and does not directly enter DPD. Nevertheless, it can be measured in DPD and therefore delivers useful input for the material modeling and its accuracy within the DPD approach. The same arguments hold true for thermal expansion. For a solid material, the value of the thermal conductivity increases with temperature and is normally almost independent of pressure changes. The thermal conductivity j can be calculated from the thermal diffusivity a if the density q and the specific heat capacity cp is known: j ˆ aqcp

…41†

The values of the density can be taken directly from the pvT plot by extrapolating the specific volume to ambient pressure and calculating the reciprocal value for each temperature (Fig. 3-14). The measurement of the specific heat capacity is typically carried out by differential scanning calorimetry (DSC) using a sapphire sample as reference (Fig. 3-15). The thermal diffusivity can be measured precisely with the laser-flash method allowing for short measurement times even at higher temperatures. The heating source and the temperature sensor are coupled without contact, avoiding extensive preparation of the sample. A short laser flash heats the top surface of the discoidal specimen (ca. 1 mm thickness). Additionally, the whole assembly can be installed in a furnace to measure the values of a at higher temperatures. The inserted heat spreads over the specimen and warms the lower surface. The time-dependent temperature distribution is given by the thermal diffusivity a. An infrared sensor detects the warming and the measured signal is plotted. Mathematical analysis using numerical models provides the thermal diffusivity and subsequently the thermal conductivity can be calculated. However, since the thermal conductivity must be experimentally determined by measuring the thermal diffusivity, the mass den-

3.4 Matching Macroscopic Parameters with DPD

Fig. 3-14 Calculated thermal conductivity and measured specific heat capacity of a l-PIM feedstock material filled with 174 PH powder (94 wt%) in the relevant molding temperature range

sity and the specific heat capacity as given in Equation (41), it might be useful to reproduce these quantities by directly DPD. In order to measure thermal expansion, a specimen is prepared and placed at the bottom of the outer dilatometer tube with the inner one resting on the specimen. The digital indicator, firmly attached to the outer tube, is in contact with

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5.6

density [g · cm3]

78

5.5

5.4

5.3

5.2 20

40

60

80

100

120

140

160

temperature [8C] Fig. 3-15 Temperature-dependent density change calculated from pvT measurement of a l-PIM Zirconia feedstock

the top of the inner die and indicates variations in the length of the specimen with changes in temperature (see Fig. 3-16). Temperature changes are efferted by a furnace in which the complete experiment is performed. The thermal expansion coefficient a is defined as [61] aˆ

DL LDT

…42†

where L is the initial length of the specimen, DT is the change in temperature and DL is the change in length. Measurements are performed on specimens cut from molded parts in the direction of flow and transverse to it (Fig. 3-17). The representation of the temperature as an internal variable of the DPD particles requires the introduction of an additional degree of freedom. Using Español’s notation [25], this degree of freedom ei is introduced for each particle, representing its internal energy, and, additionally, an entropy si ˆ s…ei † which is needed for the definition of a ‘temperature’ Ti = (@si =@ei † 1 for each particle.

Fig. 3-16 Tube dilatometer

3.4 Matching Macroscopic Parameters with DPD

The additional equation of motion for ei is e_ i ˆ

X m _D vij
F D ij ij ‡ q 2 j;j6ˆi

m 2 2 1 X m rij wR …rij † ‡ p vij
F Rij ‡ q_ Rij 2 Dt j;j6ˆi 2

…43†

In [24], q_ D ij is called ‘mesoscopic heat flow’ and is computed as follows:  q_ D ij ˆ jij

1 Ti

 1 we …rij † Tj D

…44†

Additionally, there is a ‘random heat flux’ [25] q_ Rij with q_ Rij ˆ aij wRe …rij †neij

…45†

jij can be interpreted as the thermal conductivity between two particles. It depends on the particle energies of individual particle pairs and it is assumed that jij = jji . The factor aij is a noise amplitude. Also, c and r from Equations (5) and (6) now transform, in principle, to coefficients cij and rij for individual pairs. wDe (rij ) and wRe (rij † are additional weight functions needed for the corresponding changes of the particle energies. neij is a second random number with the same characteristics as for nij in Section 3.3. For simplicity, it is assumed

Fig. 3-17 Measured linear expansion coefficient of a commercially available zirconia feedstock calculated from experimentally determined thermal expansion data

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here that all particles possess the same mass m. Otherwise, for the particle interactions, a geometric mean would have to be calculated. Again, the coefficients are not independent, but the following relations hold: wRe 2 ˆ wDe

and

a2ij ˆ 2jij

…46†

Additionally, Equation (7) still holds and Equation (8) is replaced by m r2 2kB Tij ij

…47†

  1 1 1 ˆ ‡ 2 Ti Tj

…48†

cij ˆ with Tij

1

representing a mean inverse temperature of two particles i and j. The thermal conductivity as measured by experiment (Fig. 3-14) must be reproduced by the coefficients entering Equations (44) and (45). Therefore, the respective measurement of thermal conductivity has to be performed in DPD. We expect that the Lowe–Andersen approach (see Section 3.4.3) is also applicable in this case. 3.5 Test Geometries

To verify our computational approach, test geometries for experimental verification were developed. All molded parts consist of two identical specimens that are connected in one layer rotated at 1808. By placing the sprue in the middle of the shape, identical component geometries along the melt flow path are obtained. Multiple temperature and pressure sensors are positioned in the cavity to characterize the state of the feedstock related to the position and to compare it with simulation results. The different mold inserts serve for the evaluation of the behavior of the feedstock. The test geometries were performed in double cavities to realize the implementation of the instruments (Fig. 3-18). The mold inserts shown serve as a compromise between minimal sensor diameter and real microcavities. With respect to the sensor size (diameter 2.5 mm for pressure sensor) and the required supplies it is currently not possible to create smaller cavities without affecting the material flow by the sensors. The different test geometries are designed for different purposes as follows: · Spiral: for investigation of the bulk factor in long cavities, i.e. ratio of flow length to wall thickness. · Direction change: angle of 908 to investigate the temperature and pressure drop and, as necessary, segregation at the corner.

3.5 Test Geometries

Fig. 3-18 CAD drawings of the specimen in the test geometries tool. For a description of the geometry application, see the text. The photograph shows the mold insert for the ‘spiral’ cavity with the connectors for heating and cooling

· Injector: component with an obstacle behind the injection point to analyze the material blending or transport connection after the barrier, characterization of compression effects. · Diffusor: component with an obstacle at the end of the material flow to study the filler content and the flow line problem and also characterization of decompression effects. · Stair: part with steps to analyze the temperature and pressure changes in the mold insert before and after steps. · Bending bar: the green compact and the sintered CIM part can be used directly for the determination of solid-phase material properties. Simulations with c-CFD programs have shown that most of the test cases cannot be computed precisely enough to give valuable predictions for the process parameters needed in Micro PIM [65]. Moreover, it is very difficult to incorporate models for the description of particle flow in a continuum approach, whereas in DPD the problem of a flow of solid particles in a liquid environment is inherently accounted for.

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3.6 Conclusion

This chapter has reported investigations in modeling and simulation of Micro PIM with a new fluid dynamic description using DPD. The individual parameters that enter DPD modeling of this process have to be adjusted using different specially designed experiments for the injection molding process. The most obvious difference with respect to conventional approaches is that the parameters measured in the experiment also have to be implemented as special measurement experiments in the DPD, similarly to what is known from MD. The advantage of DPD over c-CFD tools is that the computational effort for free surface dynamics is lower and the modeling flexibility is larger. This is of particular interest when two-phase flow comes into play, as is the case in Micro PIM. The DPD model cannot replace a c-CFD modeling and simulation but it can help to obtain more precise simulation results in regions where micro-features are relevant and are, by their nature, not incorporated in a c-CFD simulation program. This raises the question of proper handling of simulation regions with the respective model and therefore their coupling, which is the subject of current investigations.

3.7 Acknowledgment

Financial support by the German Research Foundation (DFG) within the collaborative research center SFB 499 Development, Production and Quality Assurance of Primary Shaped Micro Components from Metallic and Ceramic Materials is gratefully acknowledged.

3.8 References 1 M. Pidria, A. Pipino, D. Vallauri, G.

4 R. M. Nieminen, From atomistic simula-

Maizza, I. Amato, Simulation practice of powder injection molding. Adv. Eng. Mater. 2001, 3, 253. 2 H.-J. Wolf, PC-Programme für die Füllsimulation beim Spritzgießen. Z. Kunstst. 1994, 84. 3 V. V. Bilovol, L. Kowalsky, J. Dusczyk, L. Katgerman, Comparative study of different simulation programs for modeling of powder injection moulding. In: Proceedings of the European Congress on Powder Metallurgy; 2001, Vol. 44, No. 3, pp. 129– 134.

tion towards multiscale modeling of materials. J. Phys.: Condens. Matter. 2000, 14, 2859–2876. 5 M. G. Giridharan et al., Multi-disciplinary CAD system for MEMS. J. Model. Simul. Microsyst. 2001, 2, 43–50. 6 P. J. Hoogerbrugge, J. M. V. A. Koelman, Europhys. Lett. 1992, 19, 155–160. 7 M. Geier, A. Greiner, J. G. Korvink, An unconditionally stable integer lattice Boltzmann method for multi-scale simulations of polymeric solutions. In: Pro-

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ceedings of POLYTRONIC 2003, Montreux; 21–23 October 2003. P. Español, Hydrodynamics from dissipative particle dynamics. Phys. Rev. E 1995, 52, 17–34. P. Español, Dissipative particle dynamics for a harmonic chain: a first-principles derivation. Phys. Rev. E 1996, 53, 1572. C. A. Marsh, J. M. Yeomans, Dissipative particle dynamics: the equilibrium for finite time steps. Europhys. Lett. 1997, 37, 511–516. C. A. Marsh, G. Backx, M. H. Ernst, Static and dynamic properties of dissipative particle dynamics. Phys. Rev. E 1997, 56, 1676. P. Español, M. Serrano, I. Zuñiga, Coarse graining of a fluid and its relation with dissipative particle dynamics and smoothed particle dynamics. Int. J. Mod. Phys. C 1997, 8, 899–908. A. J. Masters, P. B. Warren, Kinetic theory for dissipative particle dynamics: the importance of collisions. Europhys. Lett. 1999, 48, 1–7. E. G. Flekkoy, P. V. Coveney, From molecular dynamics to dissipative particle dynamics. Phys. Rev. Lett. 1999, 83, 1775. E. G. Flekkoy, P. V. Coveney, G. De Fabritiis, Foundations of dissipative particle dynamics. Phys. Rev. E 2000, 62, 2140. R. D. Groot, P. B. Warren, Dissipative particle dynamics: bridging the gap between atomistic and mesoscopic simulation. J. Chem. Phys. 1997, 107, 4423. W. Dzwinel, D. A. Yuen, K. Boryczko, Mesoscopic modeling of colloids simulated with dissipative particle dynamics. J. Mol. Model. 2002, 8, 33–43. W. Dzwinel, D. A. Yuen, Matching macroscopic properties of binary fluids to the interactions of dissipative particle dynamics. Int. J. Mod. Phys. C 2000, 11, 1-25. P. V. Coveney, K. E. Novik, Computer simulations of domain growth and phase separation in two-dimensional immiscible fluids using dissipative particle dynamics. Phys. Rev. E 1996, 54, 5134. E. S. Boek, P. V. Coveney, H. N. W. Lekkerkerker, P. van der Schoot, Simulating the rheology of dense colloidal suspensions using dissipative particle dynamics. Phys. Rev. E 1997, 55, 3124.

21 I. Pagonabarraga, M. H. J. Hagen, D.

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Frenkel, Self consistent dissipative particle dynamics algorithm. Europhys. Lett. 1998, 42, 377–382. C. P. Lowe, An alternative approach to dissipative particle dynamics. Europhys. Lett. 1999, 47, 145–151. W. K. van Otter, J. H. R. Clarke, A new algorithm for dissipative particle dynamics:. Europhys. Lett. 2001, 53, 426– 431. J. Bonet Avalos, A. D. Mackie, Dissipative particle dynamics with energy conservation. Europhys. Lett. 1997, 40, 141–146. P. Español, Dissipative particle dynamics with energy conservation. Europhys. Lett. 1997, 40, 631–636. S. M. Willemsen, H. C. J. Hoefsloot, D. C. Visser, P. J. Hamersma, P. D. Iedema, Modelling phase change with dissipative particle dynamics using consistent boundary conditions. J. Comput. Phys. 2000, 162, 385–394. S. Halstenberg, A. Quinte, H. Eggert, C. Schön, R.-P. Peters, Einsatz numerischer Verfahren zur Modellierung und Simulation von kapillarisch getriebenen Fluiden in Mikrostrukturen. FZKa-Bericht 6423; Karlsruhe: Forschungszentrum, 2000, pp. 143–148. K. J. A. Westin, K. S. Breuer, C.-H. Choi, P. Huang, Z. Cao, B. Caswell, P. D. Richardson, M. Sibulkin, Liquid transport properties in sub-micron channel flows. In: Proc. of 2001 ASME International Mechanical Engineering Congress and Exposition, November 11–16, 2001, New York; 2001. C. A. Hieber, S. F. Shen, A finite-element/finite-different simulation of the injection-molding filling process. J. NonNewton. Fluid Mech. 1980, 7, 1–32. D. Y. C. Chan, R. G. Horn, The drainage of thin liquid films between solid surfaces. J. Chem. Phys. 1985, 83, 5311– 5324. J. N. Israelachvili, Measurement of the viscosity of liquids in very thin films. J. Colloid Interface Sci. 1986, 110, 263–271. J. N. Israelachvili, Measurements of the viscosity of thin fluid films between two surfaces with and without absorbed polymers. Colloid Polym. Sci. 1986, 264, 1060–1065.

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theory for fluids with microstructure. Int. J. Eng. Sci. 1995, 33, 2297–2308. T. Hasegawa, M. Suganuma, H. Watanabe, Anomaly of excess pressure drops of the flow through very small orifices. Phys Fluids 1997, 9, 1–3. G. R. Newkome, C. N. Moorefield, F. Vögtle, Dendrimers and Dendrons. Weinheim: Wiley-VCH, 2001. S. Granick, Y Zhu, H. Lee, Slippery questions about complex fluids flowing past solids. Nat. Mater. 2003, 2, 221–227. H. C. Lau, W. R. Schowalter, A model for adhesive failure of viscoelastic fluids during flow. J. Rheol. 1986, 30, 193–206. G. C. Georgiou, M. J. Crochet, Compressible viscous flow in slits with wall slip at the wall. J. Rheol. 1994, 38, 639–655. E. E. Rosenbaum, S. G. Hatzikiriakos, Wall slip in the capillary flow of molten polymers subject to viscous heating. AIChE J. 1997, 43, 598–608. S. G. Hatzikiriakos, J. M. Dealy, Wall slip of high density polyethylene I. Sliding plate rheometer studies. J. Rheol. 1991, 35, 497–523. D. Yao, B. Kim, J. Micromech. Microeng. 2002, 12, 604–610. N. Lion, T. C. Rohner, L. Dayon, I. L. Arnaud, E. Damoc, N. Youhnovski, Z.-Y. Wu, C. Roussel, J. Josserand, H. Jensen, J. S. Rossier, M. Przybylski, H. H. Girault, Microfluidic systems in proteomics. Electrophoresis 2003, 24, 3533–3562. P. Español, P. Warren, Europhys. Lett. 1995, 30, 191–196. M. Dijkstra, R. van Roij, J. Phys.: Condens. Matter 1998, 10, 1219–1228. I. Pagonabarraga, D. Frenkel, J. Chem. Phys. 2001, 115, 5015–5026. S. Y. Trofimov, E. L. F. Nies, M. A. J. Michels, J. Chem. Phys. 2002, 117, 9383–9394. P. B. Warren, Phys. Rev. E 2003, 68, 066702. J. E. McKinney, R. Simha, Configurational thermodynamic properties of polymer liquids and glasses. I. Poly(vinyl acetate). Macromolecules 1974, 7, 894– 901. J. Cho, I. C. Sanchez, PVT relationships and equations of state of polymers. In: Polymer Handbook, J. Brandrup, E. H. Immergut, E. A. Grulke (eds.); New York: Wiley, 1999, pp. VI/591 f.

50 S. Beret, J. M. Prausnitz, Densities of liq-

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II Tooling

Advanced Micro and Nanosystems Vol. 3. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31208-0

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4 Strategies for the Manufacture of Mold Inserts J. Fleischer, C. Buchholz, J. Kotschenreuther, Institute of Production Science (wbk), University of Karlsruhe (TH), Germany Abstract Among the most promising production technologies for producing microparts in medium to high series are replication technologies, e.g. injection molding. Replication processes are characterized by the injection of feedstock/material in negative forms, the so-called cavities. To provide these cavities, different production technologies such as cutting or material removal processes can be used. This chapter presents the state of the art of special microproduction technologies and points out strategies for fabricating micromolds and cavities.

Keywords production technology; micromachining; micromold inserts

4.1 4.2 4.3

Introduction 88 Mold Inserts 88 Cutting Processes

4.3.1 4.3.1.1 4.3.1.2 4.3.1.3 4.3.1.4 4.3.2 4.3.2.1 4.4 4.4.1 4.4.1.1 4.4.1.2 4.4.1.3 4.4.1.4 4.4.1.5

Cutting with Geometrically Defined Cutting Edges 90 Diamond Cutting Tools 90 Cemented Carbide Tools 90 Cutting of Steel as the Most Important Aim of Research 91 Microdrilling 93 Cutting with Geometrically Undefined Tool Edges 95 Microgrinding 95 Ablation Processes

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Electrodischarge Machining (EDM) Processes 98 Introduction 98 General Functioning of the Process 98 Micro-EDM Processes 99 Workpiece Materials 99 Electrode Materials and Manufacturing 100

Advanced Micro and Nanosystems Vol. 3. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31208-0

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4.4.2 4.4.2.1 4.4.2.2 4.4.2.3 4.4.2.4 4.5 4.6

Laser Beam Machining (LBM) 100 Introduction 100 General Functioning of the Process 101 Laser Beam Sources and Machining Methods 101 Laser Beam Machining/Structuring 102 Conclusions 103 References 104

4.1 Introduction

Microtechnological approaches to miniaturization are based on processes transferred from silicon technology. In particular, these are photolithography, processes of depositing thin layers, etching techniques and the cost-efficient concepts of batch manufacturing (see Chapter 8). An evolutionary approach is offered by the processes of precision engineering in which conventional methods of producing miniaturized components are developed further. An example of this is the use of new methods of ultraprecision machining for manufacturing the smallest components or structures. Mechanical processes such as cutting, drilling and grinding and material removal processes such as electrodischarge machining (EDM) or laser beam machining (LBM) are used for the small and medium batch production of miniaturized systems and microcomponents. These technologies can also be applied for macroscopic components with microstructured functional elements. LBM is especially suitable for applications demanding high precision for small-scale manufacturing while excluding conventional methods. In combination with electrochemical or electrodischarge processes or with milling, it is an ideal tool for producing microstructures [1].

4.2 Mold Inserts

Micromolds place different demands on manufacturing methods. The most commonly used methods are LIGA (lithography, electroplating and molding), micromilling, EDM and LBM. LIGA is a process for producing microstructures in a resist and then electroplating them. These structures are reproduced in a metal, which can be used independently or as a tool for molding other materials. In order to be able to find out which process is the most suitable for a certain manufacturing task, a catalogue is being built which assigns certain design features to a manufacturing process. The features that determine the manufacturing method are: · hardness of the mold material; · quantities needed;

4.2 Mold Inserts

· · · · · · · ·

smallest structure size (embossed or engraved); aspect ratio; manufacturing time; machining result (burr formation, surface quality); electrical conductivity; wall properties (steep, angled); available budget; complexity of geometry (2.5D or 3D).

Fig. 4-1 shows a micromold whose smallest feature is a 50 lm radius at the tooth tip. Hence the tool diameter needs to be 100 lm or less. The workpiece material is brass. In this case, LIGA, micromilling, EDM and LBM are possible manufacturing methods. However, milling would be the fastest process with sufficient accuracy. With a mold aspect ratio exceeding three, a milling process cannot be chosen, since tools of this length are not available. The same holds true for very hard workpiece materials exceeding a hardness of 62 HRC or if very high geometric accuracies are necessary. Smallness, aspect ratio and accuracy are the benefits of the LIGA process; costs and manufacturing time are the down side. Micromilling can only manufacture materials up to approximately 62 HRC and so far is limited to a 100 lm tool diameter and an aspect ratio of 2 in the case of a 100 lm tool. Satisfying surface roughnesses of Rz & 0.3 lm, short manufacturing times and a high degree of freedom with five-axis milling are positive aspects of this process. EDM can produce smaller structures than milling in harder materials and with very high aspect ratios. It is a rather slow process, limited to electrically conductive materials, and produces a slightly worse surface roughness than milling. LBM

Fig. 4-1 REM picture of a microgear mold (source: SFB 499)

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can achieve small feature sizes (beam diameter *10 lm), high aspect ratios in almost any material in fairly short times and, in contrast to milling and EDM, suffers no tool wear. However, surface roughness and flexibility of geometry are among the disadvantages of this process.

4.3 Cutting Processes 4.3.1 Cutting with Geometrically Defined Cutting Edges 4.3.1.1 Diamond Cutting Tools In ultraprecision machining, cutting tools of monocrystalline diamond are almost exclusively used. Diamond has a very low friction coefficient and an excellent thermal conductivity, which have a favorable effect on the cutting process. The main advantages of this cutting material, however, are its great hardness and the possibility of producing a cutting edge of almost atomic sharpness. The production of extremely sharp cutting edges belongs to the most important tasks which have to be fulfilled in microcutting. A cutting edge sharpness in the sub-micrometer range allows the production of surfaces with roughnesses of a few nanometers, providing them with an optical quality. Materials used in diamond machining are aluminum, copper, brass, nickel– silver and the so-called electroless nickel, an amorphous nickel–phosphorus alloy allowing the production of a particularly good surface and low burr formation with a comparatively high hardness. The group of ferrous products, especially steels, is not yet suitable for diamond cutting. Owing to the high process temperatures and the high affinity of carbon to iron, diamond is subject to a graphitization process leading to exorbitant tool wear. There are developments at various research facilities to solve these problems and to establish a method which allows the cutting of steel with diamond tools. At present, the approach of superimposing ultrasonic excitation on the tool movement or a preceding nitriding process of the workpiece [2] seem to be the most promising approaches.

4.3.1.2 Cemented Carbide Tools At present, the cutting process most commonly used and most extensively examined is ultraprecision turning. This process is applied for the production of molding tools made of non-ferrous metals for Fresnel lenses [3]. End milling is applied in tool and mold manufacture, and also in classical precision engineering sectors such as the watch industry. This method allows the production of complex geometries and even free-form surfaces. As in conventional production technology, milling can be regarded as the most flexible microcutting process. Single-tooth diamond disk milling cutters allow the production of grooves similar to foil turning. This process is also called

4.3 Cutting Processes

Fig. 4-2 Structures produced by fly-cutting (source: [4])

fly-cutting. In contrast to foil turning, the grooves produced can be crossed one or more times at suitable angles so that columnar or pyramidal structures are manufactured (Fig. 4-2). These structures are suitable as embossing dies and injection molding tools, e.g. for optical lattice structures with typically 1–100 lines per mm. They are used in diffraction optics and in light-guiding systems. Fly-cutting tools are commercially available with minimum diameters around 100 lm. Diamond end milling cutters are commercially available with diameters down to *300 lm. These are usually straight-flute single-tooth milling cutters. For negative structures, i.e. for the end milling of grooves, the minimum groove width is limited by the tool diameter of 300 lm.

4.3.1.3 Cutting of Steel as the Most Important Aim of Research So far, microtechnology in general, but also microcutting, has been limited to the structuring of silicon and non-ferrous metals such as electroless copper, brass and plastics [5]. In the future, it will be of great importance to be able to structure steel because of its material properties that reach from hard-brittle to soft-ductile. Besides, silicon-based molds are not as wear-resistant as steel molds when ceramic feedstock is employed. Studies on the microcutting of steel have been performed particularly in Japan and Germany since the mid-1990s, but are still within the scope of research. Wearresistant mold inserts are a necessity for the economic efficiency of the molding processes. However, the bending strength of the microstructures in the mold insert in the case of high aspect ratios can also be very important for process safety, and sometimes even for the applicability of the molding. In contrast to the ultraprecision or microcutting of non-ferrous metals, diamond cannot be used for the

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microcutting of steel. Here, cemented carbide milling cutters, widely used in conventional machining, are of interest. Owing to the single grains which appear in the form of micronotches at the cutting edge, sintered carbide tools are not suitable for obtaining optical surface qualities comparable to those achieved with diamond cutting. The comparably low price and the possibility of machining steels in

Fig. 4-3 Carbide end milling cutter by Magafor (source: wbk)

4.3 Cutting Processes

Fig. 4-4 Steel mold insert (hardness 55 HRC, surface Rz = 0.3 lm) and molded part (source: wbk)

contrast to diamond tools, however, are the reasons for using sintered carbides as cutting materials for microcutting tools. The decisive feature for the tool to be suitable for microcutting is the formation of a sufficiently sharp cutting edge. Since cemented carbide is a sintered material with tungsten carbide grains as abrasive material bonded in a relatively soft cobalt phase, the size of the tungsten carbide grains determines the achievable cutting edge rounding (Fig. 4-3). Current grain sizes range from 0.5 to 1.0 lm. Thus, cutting edge radii of a few micrometers can be realized. Carbide end milling cutters are fairly commonly used in industry and are available from several manufacturers, as coated and non-coated tools and with diameters down to 0.1 mm. Tools in experimental stages reach diameters down to 0.05 mm. Regarding microend milling of hard materials such as steel, however, special attention has to be paid to process control and the bearing of the machine tool to prevent the occurrence of stochastic tool ruptures or premature wear. Furthermore, the process requires high cutting speeds. In addition, a minimum feed per tooth is required to ensure material separation. Fig. 4-4 shows a mold insert of a wheel rim for a microvehicle which was directly milled in tool steel with a hardness 55 HRC. The surface roughness achieved is Rz = 0.3 lm. The molded part has a reflecting surface and demonstrates the capability of plastic injection molding.

4.3.1.4 Microdrilling The geometries which can be produced by drilling are limited to cylindrical surfaces. The conditions in the contact zone between the tool edge of the microdrill and the workpiece are relatively complex. As the cutting speed along the major cutting edge drops to zero in the drill center, the cutting in most tool areas does not occur at an optimum speed with relatively high cutting forces [12]. In addition, a considerable axial force is acting in the tool center caused by a chisel edge which presses the workpiece material in the area of the major cutting edges. Compared with conventional drills, microdrills have a larger ratio between the core and tool diameter in order to increase the stability (see Fig. 4-5). Generally, microdrilling tools are made of high-speed tool steel (HSS) or cemented carbide. Commercially available microdrills have aspect ratios (l/d) of

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Fig. 4-5 Drill parameters and tool coatings (source: [19–22])

*15. To increase the wear resistance of the tools and thus extend the tool life, microdrills are increasingly coated with refractory carbides or nitrides by PVD or coated with diamond layers by CVD [5, 8, 13–18].

Industrial Application Microdrilling is industrially used in the following fields: · nozzles for injection plants, valves, burners; · components for synthetic fiber production, spinning nozzles; · filters; · watch bearings and cases, bracelet links; · air bearings; · medical needles; · printed circuit boards (PCBs); · elements for fiber-optics.

4.3 Cutting Processes

In terms of quantity, microdrilling tools are mainly used in the mechanical machining of PCBs. Their trend towards miniaturization and their requirements with respect to drilling quality and production costs have accelerated the development of microdrilling tools regarding precision, tool life and diameter reduction. At present, the smallest microdrilling tools have a diameter of *15 lm. Fig. 4-6 shows industrial applications of microdrilling. The left figures show blind and through holes in PCBs. A hole in steel with a diameter of 19 lm can be seen at the top right. The bottom right picture shows a 30 lm hole in a human hair with a pushed-through 25 lm thick wire.

Research The development and optimization of specifications necessary for the drilling of microholes in various classes of materials are a main focus of research. Furthermore, theoretical and experimental examinations are carried out for optimizing the microdrilling process. 4.3.2 Cutting with Geometrically Undefined Tool Edges 4.3.2.1 Microgrinding Microgrinding allows the production of planar surfaces or grooves and of the smallest components, such as miniature shafts or drills. Since the materials to be ground are mostly semiconductor materials, glasses, ceramics or sintered carbides, diamond grinding tools are primarily used. The most comprehensive investigation results are available for the grinding processes which were developed for the production of microelectronic components, especially for monocrystalline silicon. In 1990, Tönshoff et al. published a compendium of mechanical machining steps by geometrically undefined tool edges which are used for the production of silicon wafers [25]. These include the processes of inside diameter abrasive cutting (slicing), lapping by means of blades or wire, wire grinding, surface grinding and dicing, which are necessary for the production of wafers and singles chips and have been further developed by many researchers since then [26–35]. A survey of the processes which are currently state of the art is given in Fig. 4-7. The grinding wheels used for straight grooves may have very small tool widths of *1–15 lm. A special type of grinding is the so-called dicing, by which wafers are divided into single chips using grinding wheels with a typical width smaller than 100 lm. An important field of application of this technology is the production of read/write heads for magnetic film memories for which the machining of thicker substrates of different ceramics is required. The further development of and research on dicing aim to maximize the wafer or substrate surface, which

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4 Strategies for the Manufacture of Mold Inserts Drilling in PCBs

Micro drilling in steel

Drilling a hair

according to: National Jet Company, USA

Fig. 4-6 Examples of microdrilling (source: [20, 23, 24])

is available for chips or MEMS, by means of slots that are as narrow as possible and have minimum chipping of the edges on the front and back sides of the substrate and also to increase the output by high feed speeds [37–39]. As regards the production of microstructured high-performance solar cells, the demands are clearly higher with maximum chipping of edges of *2 lm. Apart from the further development of machine system and the tools, automated quality control is of great importance. A distinction is made between dicing, in which the cut is made completely through the substrate into the carrier base, and scribing, in which the substrate is not fully separated and grooves are produced [40]. An 80 lm high and 5 lm wide and long web (aspect ratio 1 : 16) was produced by Grundig on a high-duty dicing machine using a resin-bonded dicing blade with diamond micrograin

4.3 Cutting Processes

Fig. 4-7 Survey of kinematics for microgrinding (source: [36])

sizes of 3.5–3 lm, a cutting width of 100 lm, with a cutting speed of 85 m/s and a feed of 120 mm/min (Fig. 4-8, left). The high precision of the dicing technology was also demonstrated by Disco Corporation [41] by microstructuring a human hair with a dicing blade having a width of 5 lm (Fig. 4-8, right). As the demands on the cutting width are getting ever higher, where undesirable chipping of edges occurs, ductile grinding with almost no chipping of edges in the processes of dicing or profile grinding has also attracted interest [37, 44]. In addition to microstructures which can be produced by dicing, it is also possible to produce structures using microabrasive pencils, hollow abrasive pencils (diamond hollow drills) and peripheral grinding wheels. Here, a distinction has to be made between three- and five-axis machining. Abrasive pencils are applied in the watch industry and in precision engineering. Additional application areas are, e.g., the mold, tool and model manufacture. Since 1997, conventionally designed microabrasive pencils have been available with a minimum diameter of down to 0.2 mm [45]. These microabrasive pencils can be used in various applications, which differ in the contact surfaces between the tool and the workpiece. In the case of an almost point-shaped contact, it was possible to grind free forms such as aspherical surfaces in sintered carbide with a minimal radius of *0.25 mm [46]. By means of a resin-bonded diamond abrasive pencil having a diameter of 0.25 mm and a grain size of D = 8–16 lm, a surface roughness Ra = 4 nm was produced. Peripheral grinding wheels have also been used for the production of closed microstructures. In 1999, Suzuki et al. demonstrated that it is possible to manufacture molding tools in sintered carbide using profiled microgrinding wheels by producing an example of a Fresnel lens structure [47].

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Fig. 4-8 Test piece of ferrite (left; source: [42]) and microstructured human hair (right; source: [43])

4.4 Ablation Processes 4.4.1 Electrodischarge Machining (EDM) Processes 4.4.1.1 Introduction The industrial utilization of the electrodischarge effect began in 1954. First, minimum holes, e.g. injection nozzles, were produced on cavity-sinking EDM machines. The first work concerning microelectrodischarge machining began in 1967–69 [48, 49]. The process became more interesting in the 1980s with the propagation of silicon-based microelectromechanical systems (MEMS). At present the main fields of application are holes for printer heads of inkjet printers, spinning nozzles, injection nozzles, turbine blades, electron tube grids, tools for punching of e.g. electronic components, microreactors, microtoothed wheels and mold inserts for injection molding [50, 51].

4.4.1.2 General Functioning of the Process According to DIN 8590, EDM, especially thermal removal by electric gas discharge with the spark as energy carrier, belongs to material removal processes. An electrically generated spark flashes over between the electrode and the workpiece. This discharge has a material removal effect. A dielectric between the electrode and the workpiece functions as an insulator. Another function is taking up the removed particles and conveying them out of the machining area. The sparks arising one after the other melt the electrode and the workpiece, throw parts of them into the dielectric and leave a residue of the melt re-solidifying on the material surface. The resulting roughened surface with many small craters depends on the process parameters. By adopting suitable parameters,

4.4 Ablation Processes

the removal of the undesired material from the electrode, which has the negative contour of the form to be produced, can be kept much smaller than that of the workpiece. With a multitude of removal pulses it is possible to remove a large volume of material.

4.4.1.3 Micro-EDM Processes The two most important variants of EDM are wire EDM and microcavity-sinking EDM. In the former process, a wire is used to cut out a contour from the workpiece and in the latter a cavity is produced by means of a negative form. Since all sides of the electrodes can be used as a tool, it is possible to produce a multitude of forms. For the production of complex 3D geometries, among other methods EDM is used. Here, the form is generated over several paths using mostly cylindrical rotating electrodes [52–54], gradually removing up to 100 lm thick layers from the workpiece until the final depth is reached [55]. Typical accuracies which are currently achievable in micro-EDM are presented in Table 4-1 [56, 57]. The smallest electrodes produced have a diameter of 4.3 lm at a length of 50 lm. Depending on the electrode diameter, today aspect ratios of the cavities of 10–50 are reached. Sato et al. investigated holes of 15–300 lm which found application in inkjet nozzles of printers [58].

4.4.1.4 Workpiece Materials Owing to its effective principle, the process allows workpieces to be machined independently of their hardness and strength. A certain minimum of electrical

Table 4-1 Achievable workpiece accuracies

Parameter

Wire EDM

Micro cavity-sinking EDM

Form deviation (lm) Positional deviation (lm) Median roughness, Ra (lm) Mean roughness depth, Rz (lm)

±1 <±1 0.1–0.2 0.5–2.5

±2 < ±1 0.2–0.3 1–5

Minimum structure width (lm): Webs Grooves

20–40 50–60

20–40 20–40

20–30 60–80

15–25 10–25 10–25 10 10

Maximum aspect ratio: Webs Grooves Holes Internal radius (lm) Smallest electrode dimensions (lm)

20 30

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conductivity is needed, however, in order to permit EDM. Related research studies were carried out by Reynaerts and co-workers [59–61] and Masaki et al. [62].

4.4.1.5 Electrode Materials and Manufacturing Die-sinking EDM In conventional EDM, copper and graphite are primarily used. Because of its high price, tungsten–copper is preferably used for low-wear machining, in particular for filigree structures and for the machining of carbide. Silver–tungsten and tungsten carbide are also used as electrode materials [63, 64]. For drilling up to a diameter of 0.1 mm, it is possible to use tube electrodes, allowing higher removal rates and aspect ratios to be achieved owing to the inner flushing. Electrodes with diameters below 100 lm are currently available as solid material only. The smallest available rod electrodes of sintered carbide currently have a diameter of 50 lm. If smaller electrodes are needed, these have to be self-produced. Alternative methods of producing electrodes are the LIGA process, conventional cutting, ultraprecision machining and hot pressing [63, 65].

Wire EDM Wire EDM places special demands on the electrode. Because of the running-off electrode, the wear at the wire is of secondary importance. When using thin wires (diameter 20–100 lm), attention has to be paid that the tensile strength of the wires is not exceeded. This risk can be reduced by the thermal process. Therefore, mainly tungsten wires are used which have a tensile strength of about 2500 N/mm2 [66]. Chapter 7 gives further information about the EDM process and process applications. 4.4.2 Laser Beam Machining (LBM) 4.4.2.1 Introduction The first realization of a functioning laser was successfully achieved by Maiman in the USA in 1960. This first laser set-up consisted of a rod-shaped ruby body which was excited by a flashlight [67]. Although this laser type is no longer important today, it was the impetus for the technological development of the current laser systems and applications. Development will continue to be rapid in the coming years, particularly since utilizing innovative technologies is becoming more and more important in material machining owing to increasing competitive pressures.

4.4 Ablation Processes

4.4.2.2 General Functioning of the Process LBM is based on the impact of high-energy electromagnetic radiation on a workpiece. For this purpose, the radiation is directed in a defined way on the surface of the workpiece and formed to a desired power density and distribution by optical means. The absorption of the radiation in the workpiece or at its surface results in the desired influence on the material; in other words, the energetic effect on the material occurs thermally. The laser beam ‘tool’ is not subject to wear and has high flexibility. Depending on the intensity and the qualities of the material, there are different interaction processes which are used for the design of the method. Depending on the laser beam source, there are different machining mechanisms. With CO2 and solid-state lasers, it is possible to use two different mechanisms depending on the process control. In laser beam melting machining, the material is transferred to a liquid phase through the energy input of the laser and then blown off the machining spot by means of a gas flow. This method is superior because of high removal rates for medium surface qualities. With higher beam intensities, the material is directly vaporized, resulting in finer structures with better surface accuracies, but reduced removal rates. In the above-mentioned laser types or mechanisms, the laser beam can be guided on to the workpiece surface either by a scanner or directly by an optic. In excimer lasers the material removal is generated by electron band transitions and a mask projection process is used for beam formation [67–69].

4.4.2.3 Laser Beam Sources and Machining Methods For laser beam machining, mostly pulsed laser beam sources are used, the machining being possible for the whole spectrum of materials (metals, plastics, ceramics). Common pulsed laser systems for material removal are listed in Table 4-2 [1]. Laser beam machining is applicable in the following fields: · lithography; · boring and drilling; · labeling; · engraving; · caving; · structuring; · removal of surface layers/cleaning.

Especially in tool and mold manufacture, the laser is increasingly used as a machining tool because it is very flexible and independent of the mechanical properties of the material, such as hardness or strength. Therefore, it is possible to machine materials which are difficult to cut in addition to small structures. Two main fields of laser application for tool manufacture are described in the following.

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Laser system

Pulse energy

Max. power

Repetition rate

Wavelength

Nd : YAG laser Diode-pumped qualityswitched Nd : YLF laser Diode pumped Nd : YAG laser

70 J 135 lJ

300 W 6W

1–300 Hz 1064 nm 0.1–20 ms 0.5–50 kHz 1074 nm 25 ns

0.5 W

1–500 Hz

Titanium sapphire laser

1 mJ 520 lJ 280 lJ 100 lJ 200 mJ

2W

10 Hz

TEA CO2 laser Excimer laser

0.18–50 J 10 J

5 kW 1 kW

30 kHz 1–300 Hz

1064 nm 532 nm 355 nm 266 nm 780 nm 390 nm 10.6 lm 308 nm 248 nm 193 nm 157 nm

Pulse length

30–50 ps

150 fs 6 lJs 10–30 ns

4.4.2.4 Laser Beam Machining/Structuring The first plants for LBM, in which a CO2 laser beam source was used [70, 71], were introduced at the end of the 1980s. Since the CO2 laser beam could only be guided to the machining spot via large-scale mirror systems and optics, the workpiece geometries to be machined were very limited and the plants very inflexible. In the process of laser caving, the material is removed layer by layer through the effect of the laser beam. A characteristic feature is pulsed operation with which it is possible to achieve energy densities of up to 1000 W/cm2 with very short pulse times. In the case of materials that are difficult to machine, this leads to sublimation removal without thermal effects on the adjacent material. At present, it is basically possible to produce structures down to a size of 10 lm. Depending on the laser beam sources used and the machine design, the plants that are currently on the market allow the production of structures up to a size of 100 lm with an aspect ratio of 1 : 10. The accuracy currently is 0.03 mm, which is not sufficient for micromachining. Hence there is still a need for the further development of laser beam sources for micromachining. Moreover, the surface machined by a laser needs a further finishing to remove residues of the removed material [72]. Fig. 4-9 shows a component produced by laser caving [73]. Depth engraving is a commercialized process in tool and mold manufacture which above all is used for the production of inscriptions. In this process also the power density of the laser beam is so high that the material vaporizes within a few nanoseconds during the machining. The surface and burr quality obtained by laser depth engraving can be compared with the results obtained by EDM. With steel, removal rates of more than 10 mm3/min can be reached [72].

4.5 Conclusions

Fig. 4-9 Microstructured part and typical laser caving parameters [73]: 1–5 lm material removal; Rz < 6 lm (Ra = 1.3 lm); maximum removal rate, 10 mm3/min; material, steel; laser, Q-switched Nd:YAG; maximum machining depth, 10 mm

Apart from flexibility and speed, the laser engraving process stands out with respect to zero wear when compared with conventional methods such as EDM or high-speed milling. More detailed information regarding the laser ablation process can be found in Chapter 6.

4.5 Conclusions

As described above, different production technologies can be employed for the manufacture of micromolds and cavities. The processes of cutting and material removal are suitable to meet the requirements especially for molds within replication technologies with abrasive material (e.g. powder-injection molding). In nearly all cases of manufacturing, EDM and LBM are possible processes to achieve microstructures in hard material. For structures with aspect ratios < 3 and hardness of workpiece materials < 62 HRC, cutting processes can also be used. The selection of the appropriate process technology with respect to manufacturable part features must be very specific. Therefore, a catalogue is being built up which assigns certain design features to a manufacturing process (see Chapter 18).

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4.6 References 1 R. Weichenhain, E. W. Kreutz, ‘Struktur-

2

3

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ieren mit Laserstrahlung, Veröffentlichung,’ Laser-Praxis April 1999. E. Brinksmeier, J. Dong, R. Gläbe, ‘Diamond turning of steel molds for optical applications,’ in: Euspen Proceedings, Glasgow; 2004. O. Rübenach, Schwingungsunterstützte Ultrapräzisionsbearbeitung optischer Gläser mit monokristallinen Diamantwerkzeugen; Dissertation, RWTH Aachen, 2001. http://www.ipt.fraunhofer.de, 2004. Th. Schaller, W. Bier, G. Linder, K. Schubert, ‘Mechanische Mikrostrukturierung metallischer Oberflächen,’ F&M 1994, 102, No. 5–6, 274–278. T. Moriwaki, E. Shamoto, ‘Ultraprecision diamond turning of stainless steel by applying ultrasonic vibration,’ Ann. CIRP 1991, 40/01, 559–562. D. Spath, T. Konold, ‘Mikrozerspanung – eine interessante Alternative zur Herstellung von Mikrostrukturbauteilen,’ wt Werkstattstechnik 1996, 86, 579–582. J. Schmidt, V. Hüntrup, H. Tritschler, ‘Mikrozerspante Formeinsätze aus Stahl,’ wt Werkstattstechnik 1999, 89, 495–498. A. Peichl, V. Schulze, D. Löhe, H. Tritschler, D. Spath, ‘Microcutting of steels – interaction of material properties and process parameters,’ in: Proceedings of Micro.tec 2000; 2000. F. Klocke, O. Rübenach, S. Zamel, ‘Möglichkeiten und Grenzen der spanabhebenden Fertigung bei der Herstellung von Präzisionsbauteilen,’ in: Proceedings of MicroEngineering 96, Stuttgart, 11–13 September 1996; 1996. E. Brinksmeier, ‘Herstellung von Präzisionsbauteilen durch Mikrozerspanung,’ Industriediamantenrundschau 1994, 4/94, 210–217. W. König, F. Klocke, Fertigungsverfahren Drehen, Fräsen, Bohren, 7th edn.; Berlin: Springer, 2002, pp. 375–387. R. Brück (ed.), Angewandte Mikrotechnik; Munich: Carl Hanser Verlag, 2001.

14 V. Rocek, ‘Optimierte Wendelbohrer-

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querschnitte,’ Tech. Rundsch. 1981, 73, No. 28, 7. W. Riesen, ‘Herstellen kleiner Löcher in der Feinwerktechnik (Mikromechanik),’ VDI-Ges. Produktionstechnik (ADB) 1988, October, 33–46. K. Rother, ‘Mikrobohren auf 30 Mikrometer,’ WB Jahrg. 2000, 133, No. 1–2, 60–61. Heun Werkzeugmaschinen und Industriebedarf GmbH, ‘Kleine Bohrungen bei höchster Qualität,’ DIMA 2000, No. 1–2, 16. E. Westkämper, H.-W. Hoffmeister, J. Gäbler, ‘Spanende Mikrofertigung: Flexibilität durch Schleifen, Bohren und Fräsen,’ F&M 1996, 104, No. 7–8, 52– 530. http://www.titex.com/de, 2004. http://www.ham-tools.com, 2004. http://www.rhobest.com/, 2004. H. Dimigen, J. Gäbler, ‘Potentiale superharter Werkzeugbeschichtungen aus Diamant und CBN,’ in: Lösungen aus Forschung und Industrie. 9. Internationales Braunschweiger Feinbearbeitungskolloquium (9. FBK), Oktober 1999; 1999. National Jet Company, http://www.najet.com/, 2004. HPTEC, http://www.hptec.de/, 2004. H. K. Tönshoff, W. von Schmieden, I. Inasaki, W. König, G. Spur, ‘Abrasive machining of silicon,’ Ann. CIRP 1990, 39/2, 621–635. H. Hinzen, ‘Flachschleifen von Halbleiterwafern,’ IDR 1992, 2/92, 73–77. A. Büttner, ‘Diamantwerkzeuge für die Elektronikindustrie,’ Feinwerktech. Messtech. 1985, 93, 149–152. H. K. Tönshoff, M. Hartmann, ‘Präzisionstrennschleifen von Halbleitersubstratscheiben,’ VDI-Ber. 1996, No. 1276, 383–397. H. K. Tönshoff, H. Wobker, M. Hartmann, ‘Verfahrensvergleich zum Trennschleifen von Silizium,’ IDR 1995, 1/95, 38–42.

4.6 References 30 R. Loup, ‘Sägesysteme für das Trennen

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36 37

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40 41 42 43 44

von Halbleitersubstraten mit Diamantwerkzeugen,’ IDR 1983, 3/83, 131–135. E. Brinksmeier, W. von Schmieden, ‘IDcut-off grinding of brittle materials,’ Ann. CIRP 1987, 36/1, 219–222. T. Abe, ‘A future technology for silicon wafer processing for ULSI,’ Precision Eng. 1991, 13, 251–255. E. Brinksmeier, W. von Schmieden, ‘Werkzeugaufspannung und Prozessverlauf beim ID-Trennschleifen,’ IDR 1988, 4/88, 214–219. M. Kerstan, A. Ehlert, D. Helmreich, J. Beinert, W. Döll, M. Schinker, ‘Ultraprecision Grinding and Single Point Turning of Silicon Wafers and their Characterization,’ in: Proceedings of the ASPE, Spring Topical Meeting; 1998, Vol. 17, pp. 32–35. J. Mätzig, Ultrapräzisionsbearbeitung von Strukturkeramiken mittels Nanoschleifen; Dissertation, Fortschr. Ber. VDI Reihe 2, No. 466. Düsseldorf: VDI Verlag, 1998. IWF, Braunschweig, http://www.iwf.ing.tu-bs.de/, 2004. H. Gatzen, ‘Dicing challenges in microelectronics and micro electro-mechanical systems (MEMS),’ in: Proceedings of the 3rd International Conference on MicroMaterials, MicroMat 2000, 17–19 April 2000, Berlin; 2000, pp. 141–146. H. Shah, S. Ram, ‘Characterization of the wafer dicing process using Taguchi methodology,’ in: Proceedings of the IEEE/ SEMI, ASM; 1992, pp. 200–205. J. Ikeno, Y. Tani, A. Fukutani, ‘Development of chipping-free dicing technology applying electrophoretic deposition of ultrafine abrasives,’ Ann. CIRP 1991, 40/1, 351–354. Minitron Elektronik GmbH, Company Literature; 1996, MRD-3, No. 11-96. Disco Corp., Japan, Company Literature; HD 1995. 11. 2000, 2000. Grundig AG, http://www.grundig.de. Disco Corp., http://www.disco.co.jp/. H. Gatzen, M. Beck, J. Zeadan, ‘Technology challenges in ultraprecision grinding of flexural resonators,’ in: Proceedings of the EUSPEN. 1st Int. Conf. for Precision Engineering and Nanotechnology, 31 May– 4 June 1999, Bremen; 1999, pp. 538–541.

45 R. Haefeli, ‘Anwendungsmöglichkeiten

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von Schleifstiften,’ IDR 1997, 2/97, 164– 168. H. Suzuki, O. Horiuchi, H. Shibutani, T. Higuchi, ‘Precision grinding of microaspherical surface,’ in: Precision Engineering at the Beginning of a Millennium, 10th International Conference on Precision Engineering (ICPE), Yokohama, 18–20 July 2001; 2001, pp. 361–365. H. Suzuki, T. Higuchi, T. Kitajima, S. Okuyama, N. Wajima, ‘Precision grinding of microfresnel lens molding die – feasibility study on precision grinding of tungsten carbide,’ in: Proceedings of the EUSPEN. 1st Int. Conf. for Precision Engineering and Nanotechnology, 31 May–4 June 1999, Bremen; 1999, pp. 318–321. H. Kurafuji, T. Masuzawa, ‘MicroEDM of cemented carbide alloys,’ J. Jpn. Soc. Electr.-Mach. Eng. 1968, 2, No. 3, 1–16. C. van Osenbruggen, ‘Microsparc erosion,’ Philips Tech. Tijdschr. 1969, 20, 200–213. D. Allen, A. Lecheheb, ‘Microelectro-discharge machining of ink jet nozzles: optimum selection of material and machining parameters,’ J. Mater. Process. Technol. 1996, 58, 53–66. F. Michel, W. Ehrfeld, H. Lehr, A. Wolf, H. Gruber, A. Bertholds, ‘Mikrofunkenerodieren als Strukturierungsverfahren in der Mikrotechnik,’ in: Tagungsband zum 41. Internationalen Wissenschaftlichen Kolloquium. Wandel im Maschinenbau durch Feinwerktechnik und Mikrosystemtechnik, Ilmenau, 23.–26. September 1996; Ilmenau: Universitätsbibliothek der TU Ilmenau, 1996, pp. 233–236. K. Bollen, W. van Dijck, Development of Machining Strategies for Complex (3D) Surfaces with EDM Milling; Master Thesis, K.U. Leuven, 1999. T. Kaneko, M. Tsuchya, T. Fukushima, ‘Improvement of 3D NC contouring EDM using cylindrical electrodes,’ in: Tagungsband zum 9th International Symposium for Electromachining (ISEM IX), Nagoya, Japan; 1989, pp. 49–52. J.-P. Kruth, P. Bleys, ‘Machining curvelinear surfaces by NC electro discharge machining,’ in: Proceedings of 2nd International Conference on Machining and

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Measurements of Sculptured Surfaces, Krakow, 20–22 September 2000; Krakow: Institute of Metal Cutting, 2000, pp. 271– 294. M. Shinkai, T. Suzuki, ‘An electrical-discharge scanning machine,’ in: Mitsubishi Electric Corp. Catalogue 1997; Mitsubishi, 1997, pp. 8–10. D. Allen, H. Almond, J. Bhogal, A. Green, P. Logan, X. Huang, ‘Typical metrology of microhole arrays made in stainless steel foils by two-stage microEDM,’ Ann. CIRP 1999, No. 48, 127– 130. O. Koch, F. Michel, H. P. Gruber, ‘Mikrobearbeitung durch Funkenerodieren – Technologien und Anwendungen,’ in: Seminarunterlagen zu Präzisions-, Ultrapräzisions- und Mikrobearbeitung mit Verfahren der Zerspan- und Abtragtechnik, Dresden 23 March 2001; Dresden: TU Dresden Institut für Produktionstechnik, 2001. T. Sato, T. Mizutani, K. Kawata, ‘Electrodischarge machine for microhole drilling,’ in: National Technical Report, No. 31; 1985, pp. 725–733. D. Reynaerts, W. Meeusen, H. van Brussel, ‘Machining of three-dimensional microstructures in silicon by electro-discharge machining,’ Sens. Actuators A 1998, 67, 159–165. D. Reynaerts, P. Heeren, H. van Brussel, ‘Microstructuring of silicon by electrodischarge machining (EDM) – Part I: theory,’ Sens. Actuators A 1997, 60, 212– 218. P. Heeren, D. Reynaerts, H. van Brussel, C. Beuret, O. Larsson, A. Bertholds, ‘Microstructuring of silicon by electro-discharge machining (EDM) – Part II: applications,’ Sens. Actuators A 1997, 61, 379–386.

62 T. Masaki, K. Kawata, T. Masuzawa, ‘Mi-

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croelectro discharge machining and its applications,’ in: Proc. IEEE Micro Electro Mechanical Systems, Napa Valley, USA, 11–14 February 1990; 1990, pp. 21–26. J. Stampfl, R. Leitgeb, Y. Cheng, F. Prinz, ‘Electro-discharge mesoscopic parts with electroplated copper and hotpressed silver tungsten electrodes,’ J. Micromech. Microeng. 2000, No. 10, 1–6. M. Her, F. Wenig, ‘Microhole machining of copper using the electrodischarge machining process with a tungsten carbide electrode compared with a copper electrode,’ Int. J. Adv. Manuf. Technol. 2001, 17, 715–719. G. Benavides, P. Adams, P. Yang, ‘Mesomachining capabilities,’ in: Sandia Report SAND2001-1708; 2001, pp. 25–28. T. Nöthe, Funkenerosive Mikrobearbeitung von Stahl und Hartmetall durch Schneiden mit dünnen Drähten; Dissertation, RWTH Aachen, 2001. D. Förster, W. Müller, Laser in der Materialbearbeitung; Munich: Carl Hanser, 2001, pp. 9–10. J. Eichler, Laser, Bauformen, Strahlführung, Anwendungen, 4. Auflage; Berlin: Springer, 2001, pp. 371–380. W. König, F. Klocke, Fertigungsverfahren, Abtragen und Generieren, 3. Auflage; Berlin: Springer, 1997, pp. 179–182. ‘Das MAHO Lasercaving – ungeahnte Dimensionen im Formenbau,’ Opto Elektronik Mag., 1990, 6, No. 1, 12–15. G. Eberl, et al. ’Neue Entwicklungen beim Laserabtragen,’ Laser Optoelektron. 1992, No. 4, pp. 44–49. Baasel Lasertechnik, Neue Laser im Formenbau; Veröffentlichung Lasermarkt, 2001. Lang GmbH, Huettenberg, www.lang.de

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5 Micro End Milling of Hardened Steel J. Schmidt, J. Kotschenreuther, Institute of Production Science (wbk), University of Karlsruhe (TH), Germany Abstract Miniaturization of devices and components is one of the main trends in technology. Microfluidic, micromechanic, microoptical and microelectronic devices are combined with microsystems, leading to new products. For a major breakthrough in microtechnology, economical processing is of vital importance. At present, most manufacturing processes are adapted from the semiconductor industry, which is economical only for large lots. Furthermore, these usually lithographic processes strongly restrict the design geometry. Microcutting has the potential to overcome these problems and to offer new fields of application to microtechnology. New developments of microcutting of steel are presented here in general in addition to research results obtained by Project B2 ‘Microcutting’ of the Collaborative Research Center 499.

Keywords micromilling; micromolds; alternative tool production methods; hardened steel

5.1 5.2 5.2.1 5.2.1.1 5.2.1.2 5.2.1.3 5.2.1.4 5.2.2 5.2.3 5.2.3.1 5.2.3.2 5.2.3.3 5.2.3.4

Introduction 108 Micromilling 111 Tool 111 Tool Materials 111

Manufacturing Methods 114 Alternative Manufacturing Methods 114 Simulation 119 Work Piece 122 Process 122 Cutting Tools 123 Theoretical Considerations for the Cutting Force 123 Measurements of the Cutting Forces 123 Theoretical Considerations for Tool Deflection 123

Advanced Micro and Nanosystems Vol. 3. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31208-0

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5.2.3.5 5.2.3.6 5.2.3.7 5.2.3.8

Surface Quality 123 Burr Formation 125 Process capability 125 Manufacturing of Mold, Models and Test Structures for Demonstrator Components and Wear Analyses 127

5.3 5.4

Conclusions 127 References 128

5.1 Introduction

A major development trend in industrial technology is the miniaturization of units and components. Microfluid, micromechanical, microoptical and microelectronic functionalities are combined on the smallest construction areas with microsystems, leading to new products in bio- and medical technology, in information and communication, in automotive engineering and in several other areas of application. For a breakthrough in microtechnology, an economic production technology is crucial [1, 2]. At present, industrially assigned manufacturing methods are usually adapted from the semiconductor industry and are therefore economical only for large lots. Additionally, the machinable geometry and structurable materials are strongly restricted when using lithographic processes. To overcome these restrictions, manufacturing process chains which are composed of classical and precision-engineering procedures are wel-suited. Molding and transforming processes offer great potential and, therefore, are intensively used nowadays for the production of microunits [3]. For the successful employment of molding and transforming processes, knowhow in molds and dies is all-important [4]. Trend-setting applications, for example in bio- and medical technology, in automotive engineering and in the optical industry reveal extreme requirements concerning surface quality and geometrical accuracy for the manufacture of micromolds [5]. At the same time, economical production requires a long tool life of the molds, which can only be achieved with wear-resistant and hence difficult to machine materials [6]. Mold production by means of electrodischarge machining (EDM) or laser ablation is problematic, since the highest surface qualities can only be achieved by extensive manual polishing. The form deviations are thereby worsened to values which are not tolerable for highly precise machining [7–9]. For such requirements, microcutting is often the only practicable solution. However, present applications of CNC (Computerized Numerical Control) milling are limited to geometry sizes starting from *100 lm, e.g. milling of mold inserts for hot embossing of microfluid analysis components [10]. Smaller structures still have to be manufactured by complex and time-consuming lithographic procedures [11]. Microinjection molding as the main scope of application for microcutting imposes high demands on the surface quality – surface roughness Rz = 0.5 lm is

5.1 Introduction

desirable – and does not tolerate burr formation. Whereas steel molds with straight geometries can be obtained rather conveniently using wire EDM, milling is the ideal machining technology for free form surfaces [12–14]. The further development of cutting processes offers great potential for the miniaturization of components regarding the cost-efficient manufacture of microcomponents in medium batch sizes. The special merits of cutting are material variety, geometric freedom, surface quality and short machining times. Concerning cutting with a very small chip cross-section, a distinction is made between ultra-precision cutting and microcutting. Ultra-precision cutting is machining of large-surface components with an accuracy in the sub-micrometer range and with optical surface qualities. Roughnesses of Ra = 0.0067 lm are possible by grinding and polishing. Microcutting refers to manufacturing by cutting of structures with micrometer dimensions. In microcutting, good results can be obtained with aspect ratios of up to 5 by using end-milling tools. Using profiled milling tools, aspect ratios as high as 10 may be achieved. Surface roughnesses of Rz = 0.1 lm may be reached by using cemented carbide tools. Depending on the workpiece material, web thicknesses of 20–50-lm and structure heights of 1-lm to several millimeters can be achieved. In contrast to the limited structure variety of ultra-precision machining, microcutting has the potential to machine free form surfaces by employing five-axis cutting. Starting from the ultra-precision cutting of non-ferrous metals with diamond, a main development trend is the structuring of super-hard, wear-resistant materials, allowing for the development of totally new applications for microsystems technology. The first tests concerning the suitability of microcutting processes were carried out in the USA at the end of the 1960s and were focused on the manufacture of surfaces of optical quality to be used in the optical industry [15]. This development led to ultra-precision machining of optical, electronic and mechanical components with the accuracies mentioned above [16]. The workpieces may have a diameter of > 0.5 m. The cutting material which is nearly exclusively used is monocrystalline diamond. At the end of the 1980s, the Forschungszentrum Karlsruhe (Karlsruhe Research Center) recognized that diamond cutting can be used not only for the production of the highest surface qualities, but also for the microstructuring of elements for microsystems technology. A first application was the production of micro heat exchanger elements [17]. For this purpose, a copper or aluminum foil was stretched on a rotating drum and provided with grooves using a monocrystalline diamond cutter. These grooves differed depending on the tool geometry. When stacking up square segments of these foils, alternately displacing them by 908 and connecting them by, e.g., diffusion welding, heat-exchange modules were obtained with a very high heat transmission coefficient and very small dimensions. Further, the ultra-precision and microcutting of non-ferrous metals has become established especially for applications in the optical industry. Applications which are used commercially with successinclude air bearings, molding tools for reflection foils and automobile lighting elements, non-spheri-

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cal copper mirrors for high-performance lasers and components for minimal invasive surgery. Until the mid-1990s, microcutting was focused on the structuring of non-ferrous metals using diamond tools [18–25]. The permanently broadening field of applications in microsystems technology, however, has made it clear that a greater variety of materials is required. Therefore, the microcutting of steel and ceramics, which is still performed within the scope of research, has recently become of incraesing interest for industry. Cutting of ferrous metals with diamonds leads to substantial wear due to the high affinity of iron to carbon [26]. At cutting temperatures above 700 8C diamond begins to degenerate to graphite, which is the stable phase of carbon at ambient temperature. Considerable research effort has been dedicated to this problem over the past few years. Three approaches may satisfy the industrial demands for cutting in hard steel: 1. Reduction of the contact time between the diamond tool and the workpiece by applying ultrasonic vibrations to the tool [22]. Fig. 5-1 shows an experimental setup of an ultrasonic-assisted diamond turning process at Fraunhofer IPT, Aachen, Germany. 2. Application of a nitriding process to the workpiece. With this thermochemical modification, chemical reactions between the carbon atoms of the diamond lattice and the iron atoms of the workpiece material can be suppressed, thus preventing chemical wear and reducing the total wear to the same magnitude as in cutting brass [27]. 3. Use of cutting materials other than diamond such as cemented carbide and cobalt-based hard metal, ceramics or monocrystalline cutting materials [28]. For end milling as the most flexible cutting process in terms of geometry, only a suitable variation of the cutting material is a promising option, since the application of ultrasonic vibrations is feasible only in turning operations. Tools with diameters > 0.1 mm are commercially available from a variety of companies. In experimental stages, tools with diameters of 0.06 mm can be ob-

Fig. 5-1 Ultrasonic-assisted diamond turning, Fraunhofer IPT

5.2 Micromilling

tained from only very few manufacturers. These tools are typically manufactured on highly precise grinding machines using grinding wheels down to 0.025 mm in diameter. A surface quality of the workpiece of Rz = 0.3 lm can be achieved with these tools. Removal rates range from 0.3 to 10 mm3/min and typically lie between 2 and 5 mm3/min. A barrier to further miniaturization of tools is their manufacture by a grinding process itself. Not only is extensive knowhow required, but also process forces inhibit the grinding of much smaller diameters. Owing to the absence of process forces, tools with diameters of £ 0.05 mm are manufacturable using ion beams or laser beams. However, up to now a geometry check of tools with these diameters is very challenging.

5.2 Micromilling

The goal of the SFB 499 subproject B2 ‘Microcutting’ is to make controllable the cutting manufacture of microstructures in hardened steel for tools and molds with three-dimensional free forming geometries and to exceed the machining results currently attainable with commercial tools. Criteria for the achievement of the objectives are the smallest possible structure size, geometric accuracy, condition of the produced edge layer and tool life. The tool and process development for microcutting is carried out in a close network with the subprojects B4 ‘Micro EDM’, B5 ‘Optimization of molds’ and B6 ‘Combination of EDM and laser ablation in microstructuring’. Together with these subprojects, a contribution to the optimization of micromold manufacture is aimed for. At the Institute of Production Science, Karlsruhe (wbk), microcutting experiments are carried out on a conventional micromilling machine built in-house and on a tailor-made, highly dynamic, high-precision micromilling machine equipped with linear motors and hydrostatic bearings. In these experiments both commercially available and tailor-made cemented carbide end milling tools are used. 5.2.1 Tool

One of the main limiting factors for the miniaturization of the cutting process is the tool. Various partially conflicting requirements need to be met in order to produce these precision components. The requirements are to be met in different fields: tool material and manufacturing method.

5.2.1.1 Tool Materials Since both surface quality and cutting forces depend strongly on the microgeometry of the cutting edge, care has to be taken with its formation. An extremely

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sharp and at the same time durable edge is desirable. Depending on the grain size, cemented carbide tools are limited in this respect to a minimum cutting edge roundness of 2–5 lm. Exceeding this limitation is only feasible by variation of the tool material. As single-crystal cubic boron nitride is hardly available in grain sizes suitable for single-point cutting tools, artificial sapphire was used to obtain a sharper cutting edge for the experiments. Judging from SEM images, the edge appears comparable to single-crystal diamond tools. It is important to utilize a tool substrate that meets the following requirements: · Ultra-fine grains (*0.5 lm) to form a sharp cutting edge. · Homogeneous grain size, to avoid tool ruptures at larger grains. · High Young’s modulus, in order to prevent tool drift. · High hardness, to resist abrasion and high temperatures. · High compression strength, to avoid plastic deformations at high temperatures. · High bending strength, to resist tool rupture due to a bending force. · High fracture toughness, to offer good resistance against crack propagation. · High resistance to fatigue due to discontinuous cut and non-uniform chip flow. · High resistance against thermal shock due to high temperatures and high temperature gradients. Furthermore, wear-reducing coatings appear essential for an economically reasonable tool wear life. Here, PVD (Physical Vapor Deposition) coating, e.g. TiAlN, but also gradient coatings or multilayer coatings of TiN yield good results [29]. The often-cited rounding of the cutting edge due to coating is almost negligible, as can be seen in Fig. 5-2. The use of a coated cemented carbide tool, however, needs dedicated process parameters, e.g. the cutting velocity must be high enough to avoid conditions resulting in brittle fracture (Fig. 5-2), otherwise

Fig. 5-2 TiAlN-coated cemented carbide tool (diameter 0.3 mm) operated at low cutting velocity vc displays coating chunking of comparatively large area

5.2 Micromilling Fig. 5-3 TiAlN-coated cemented carbide tool (diameter 0.3 mm) with droplets displays the current limits of coating technology for micromilling tools

larger areas of the coating tend to break away, whereas at cutting velocities of > 100 m/min the tool wear is of a continuous type. The formation of droplets is still a major problem even for sophisticated coating equipment for TiAlN (Fig. 5-3). The coating usually consists of a sub-micrometer TiAlN layer. A coating needs to fulfil the following criteria: · Reduce mechanical, chemical and thermal influences. · Adhere well on the substrate. · Feature a very smooth surface, no droplets. · Offer a very small thickness to keep the cutting edge sharp. Using tools fulfilling these characteristics and highly accurate, stiff and dynamic machining centers, machining of steel with a hardness of > 60 HRC becomes possible. An alternative cutting tool material is monocrystalline Al2O3, artificial sapphire, as mentioned above. Experiments with continuous cut have been carried out already showed varying results [30]. The material is very hard and may possess cutting edge radii of atomic sharpness. The tools investigated had cutting edge radii of 1–2 lm. However, this material is also very brittle. Fig. 5-4 shows the sharp cutting edge of an Al2O3 inlay in an indexable insert (left) and a broken tool after an *500-mm continuous cutting path in Armco-Iron, a model material with a low content of carbon (middle and right). A major problem is the reproducible manufacture of these tools. The delivery condition of each cutting tool varies widely so that each cutting tool has to be analyzed and checked as suitable for the process. With this kind of presupposition and brittle wear and non-reproducible tools, the application of this material for tools seems hardly possible at this time.

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Fig. 5-4 Sapphire cutting tool: sharp cutting edge, unused (left), REM picture of severed cutting tool (middle) and confocal whitelight microscope picture of dismantled tool (right)

5.2.1.2 Manufacturing Methods The common manufacturing method for cemented carbide cutting tools is grinding using diamond grinding blades with grinding wheel sizes ranging from several centimeters down to fractions of millimeters. Since grinding wheels equipped with micrometer-sized diamond grains are available, fine structuring of tool geometries down to 0.1 mm was employed fairly successfully. However, as a result of the grinding forces and vibrations, the defective rate is unacceptably high. In order to be able to produce even smaller diameters and tailored tool geometries, alternatives to the conventional manufacturing techniques were analyzed according to their potential to create cemented carbide cutting tools without process forces and vibrations. Examples of tools smaller than 0.1 mm in diameter can be seen in Fig. 5-5.

5.2.1.3 Alternative Manufacturing Methods Ion Beaming Ion beaming as an alternative manufacturing approach for milling tools has been carried out in cooperation with the Rossendorf Research Institute. Tools

Fig. 5-5 Cemented carbide milling tool with a diameter of 0.06 mm manufactured by grinding

5.2 Micromilling Table 5-1 Data for first tool manufactured by focused ion beam machining

Material Diameter Rake angle Flank angle Manufacturing time

Ultra-fine cemented carbide 0.052 mm –558 308 12 h

Fig. 5-6 Milling tool made from cemented carbide with two cutting edges manufactured by ion beaming (left), and cutting result occurring when using this tool (right)

with two cutting edges were machined with Au ions at 35 keV, a current density of 1.2 A/cm2 and a spot size of *0.3 lm. Table 5-1 shows the manufacturing data corresponding to the first tool produced by focused ion beam machining [31]. Fig. 5-6 shows that ion beam machining of cemented carbide is basically possible. However, the cutting result occurring when using a tool produced by this process is unsatisfactory regarding burr formation. The resulting surface roughness of the milled groove was *1.0 lm while achieving a milled path of *5 mm before tool rupture. To overcome this problem, the tool geometry had to be changed. As was established earlier [32], single flute end milling tools appear more reasonable for very small tool diameters owing to high radial geometric inaccuracies, unavoidable by the grinding process, which can be found on commercial tools with two flutes and lead to different chip cross-sections for each cutting edge. Therefore, both by focused ion beam (FIB) and laser machining, single flute end mills were manufactured. Table 5-2 gives the resulting geometry and the machining time, and the new geometry is shown in Fig. 5-7. Use of this tool resulted in slightly less burr formation whereas the surface roughness and tool life could not be enhanced. This may be due to deflections of this filigree tool while milling in X38CrMoV5-1 (1.2324) with a hardness of 52 HRC. Series production of milling tools by ion beaming is inefficient owing to the extremely long manufacturing time of several hours for one tool.

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5 Micro End Milling of Hardened Steel Table 5-2 Data for second tool manufactured by focused ion beam machining

Material Diameter Rake angle Face angle Manufacturing time

Ultra-fine cemented carbide 0.047 mm 08 108 *20 h

Fig. 5-7 New geometry with single flute manufactured by focused ion beam machining

Laser Ablation Another manufacturing method is laser ablation. The experiments were conducted at LaserPluss AG, Kirschweiler, Germany [31]. The tool was made of ultra-fine grain cemented carbide and was machined using an Nd : YAG UV laser source of 1 W at 266 nm. Table 5-3 shows the geometry and manufacturing data for the laser-machined tool. Fig. 5-8 shows the manufactured milling tool. Milling experiments with this tool were conducted in X38CrMoV5-1 (1.2324) with a corresponding hardness of 52 HRC. After *30 mm, tool rupture occurred with the cutting results displayed in Fig. 5-9. The surface roughness Rz achieved is 1.241 lm. Optimization of the milling results focuses on burr reduction, which will be pursued by appli-

Table 5-3 Data for tool machined by laser ablation

Material Diameter Rake angle Face angle Manufacturing time

Ultra-fine cemented carbide 0.2 mm 08 98 *30 min

5.2 Micromilling

Fig. 5-8 Laser-machined carbide tool (diameter 0.2 mm) (middle and right), compared with a similar geometry machined by grinding (upper and lower left)

cation of a higher rake angle and eventually by application of a sharper cutting edge using a different tool material.

EDM The last manufacturing method tested in this context was electrodischarge manufacturing (EDM). These experiments were conducted at the Institute of Production Science, University of Karlsruhe (TH) [33]. Further information regarding the manufacturing process of these tools can be seen in Chapter 7. Owing to inaccurate clamping, the shaft of the tool had to be manufactured by electro-

Fig. 5-9 Milling of steel X38CrMoV5-1 using a laser-ablated cemented carbide tool: tool path in workpiece (left) and bottom of groove with clearly visible feed marks (lower right). Picture taken by optical profilometry

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5 Micro End Milling of Hardened Steel Table 5-4 Data for tool produced by EDM

Material Diameter Rake angle Face angle Manufacturing time

Ultra-fine cemented carbide 0.5 mm 08 108 60 min

discharge grinding on the same machine. This step took about 5 h for each tool. With highly precise chucking now available, this step can be omitted, thus reducing the manufacturing time to *1 h. Two different tool materials were used in the course of the experiments, mainly varying in mean carbide grain size (0.8 and 0.4 lm) and hardness (HV30 of 1730 and HV30 of 1560) to discover correlations between tool, manufacturing process and milling process. Ten tools of each raw material were manufactured. Table 5-4 shows the geometric parameters of the manufactured tools shown in Fig. 5-10. Using the tools manufactured by EDM in workpieces made of the steel X38CrMoV5-1 (1.2343), differences in the raw material of the tool became obvious. The lifetime of tools made of cemented carbide with a mean grain size of 0.4 lm and a hardness of 1730 HV30 was on average twice as long as those made of cemented carbide with a mean grain size of 0.8 lm and a hardness of 1560 HV30. Fig. 5-11 shows considerable wear after cutting a groove 20 mm in length. This may be due to a poor cutting geometry, unsuitable surface manipulation by the EDM process or the lack of a coating. Fig. 5-12 displays two milled geometries, a groove (left) of 20 mm and a partial view of a test structure (right) with a depth of 200 lm. A direct comparison with the other manufacturing alternatives in terms of tool life seems inadequate since this tool has a much larg-

Fig. 5-10 Tool made from cemented carbide manufactured by micro-EDM: general view (left) and close up (right) (mean carbide grain size 0.8 lm)

5.2 Micromilling

Fig. 5-11 Same tool as in Fig. 5-10 after 20 mm of cutting in X38CrMoV5-1 (1.2343) (mean carbide grain size 0.8 lm)

Fig. 5-12 Groove milled by cutting tool *20 mm (left), part of test structure milled with cutting tool manufactured by EDM. Cutting path: 544 mm (right). Both geometries were manufactured in X38CrMoV5-1

er diameter. The objective of further research is to miniaturize these tools and thus make a direct comparison possible.

5.2.1.4 Simulation The prediction of tool failure is supported by elastostatic calculations of tool stability, partly using FEM. In a first step, a simple geometry is used since it is accessible to rather simple theoretical considerations. The von Mises stress distribution is not applicable to ceramic or carbide tools owing to different failure criteria. However, comparisons between maximum principal stresses and von Mises stress distributions have shown very good agreement. Therefore, the von Mises stress given below is for comparison purposes only.

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5 Micro End Milling of Hardened Steel Fig. 5-13 Model of microcutting tool with stress distribution

For example, a cemented carbide tool with a diameter of 0.3 mm and a single flute cutting tip with a Young’s modulus of 549 000 N/mm2, Poisson’s ratio of 0.23 and yield strength of 1200 N/mm2 could be calculated to sustain 1.8 N of allowable cutting force (Fig. 5-13). This matches experimental tool yield observations [34]. In a further step, more complex geometries were generated. The software used was created by Rollomatic, one of the leading manufacturers of grinding machines. By entering geometry information about the tool, e.g. diameter, rake angle and coil angle and also the grinding parameters and tools, the software calculates the result of that grinding process. The extracted 3D models were implemented in the simulation software Abaqus CAE 6.4. Fig. 5-14 shows two major areas of stress accumulation: · the cutting edge which is primarily exposed to the force, and · the transition from the cylindrical shaft to the cone due to the notch effect.

Fig. 5-14 Model of tool created by the Rollomatic software in Abaqus CAE 6.4 showing the maximum principal stress distribution (static load)

5.2 Micromilling Fig. 5-15 Close up of the cutting area of the simulated milling process (static load)

While absolute stress data show a much higher value at the cutting edge (by at least one order of magnitude), most tool ruptures occur near the transition area. Fig. 5-15 shows a close-up and the meshing of the model displayed in Fig. 5-14. It can be seen that the mesh becomes finer in the vicinity of the applied force. Starting from a standard tool geometry of a 0.1-mm milling tool, several variations of a single geometric feature have been analyzed so far while maintaining all other features. The varied parameters and the effect on the normalized maximum von Mises stress can be extracted from the Table 5-5, where 1 indicates the features of the standard tool.

Table 5-5 Geometry variations with effect on normalized maximum von Mises stress (static)

Parameter

Variation

Normalized maximum von Mises stress

Minor diameter

0.03 mm 0.05 mm 0.07 mm 38 48 68 78 258 278 308 0.02 mm 0.035 mm 0.05 mm Identical with core diameter 0.03 mm

0.891 1 0.894 0.907 1 0.932 0.923 0.881 0.83 1 1 0.936 0.952 0.621

Rake angle

Coil angle

Cylindrical chamfer width

Torus milling tool

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5 Micro End Milling of Hardened Steel Table 5-6 Effects of normalized von Mises stress while changing the depth of cut and the feed per tooth

ap (mm)

fz (mm)

Normalized resulting max. von Mises stress

0.01 0.01 0.02 0.02

0.002 0.004 0.002 0.004

1 1.219 1.001 1.208

All variations are within the same order of magnitude. Only slight reductions of the maximum stress can be observed. To minimize the forces that act on the corner of the cutting edge, torus milling tools may be favored and therefore were simulated. They showed smaller maximum stresses than square end mills. In a next step, the variants with decreased stresses are combined in a new tool which will be modeled and exposed to the same load. Apart from varying the geometry, cutting conditions have been varied using the same cutting tool. The results in Table 5-6 show a comparatively small influence of the depth of cut compared with the feed per cutting edge. Therefore, to decrease the stress of a milling tool, a small decrease in the feed per cutting edge is more effective than a large change in the depth of cut. After conducting these static simulations, a calculation for the dynamic load case is in progress to fit the simulations better into the real cutting process. 5.2.2 Work Piece

The selection and conditioning of the workpiece material plays a major role in the achievement of a sufficient machining result and is crucial for the stability of the process. The current optimum is represented by vacuum-remelted or powder metallurgical steels with carbide particles < 2 lm in diameter, thus not exceeding the typical depth of cut. Two further promising alternatives regarding wear resistance while replicating with ceramics and machinability are currently being tested. These are the steels Toolox 44 and MPM 9, mostly employed within the polymer processing industry. For further information, see Chapters 9 and 21. 5.2.3 Process

Research concerning the cutting was carried out on heat-treated tool steel Toolox 44 and MPM 9 with hardnesses ranging from 42 to 52 HRC. Grooves were produced in an end milling process while measuring process forces.

5.2 Micromilling

5.2.3.1 Cutting Tools The milling tools employed are cylindrical end milling tools with center cut and diameters between 0.1 and 0.5 mm made of ultra-fine grain cemented carbide and feature a mean grain size of 0.7 lm.

5.2.3.2 Theoretical Considerations for the Cutting Force The cutting forces constrain both the tool life by rupture and the geometric accuracy due to the deflection of the cutting tool. In combination with the cutting force measurement executed as preparatory work, theoretical calculations have been considered. Following the cutting force equation of Victor and Kienzle [36]. cutting forces between *300 mN and several N are assessable.

5.2.3.3 Measurements of the Cutting Forces Cutting forces using a sharp tool (0-mm cutting path) were determined and entered into the diagrams shown in Fig. 5-16. In comparison with the estimations in the previous section, these values are slightly higher, which can first be attributed to the correction values of the cutting force calculations and second be related to the dynamic transfer function of the force-measuring platform.

5.2.3.4 Theoretical Considerations for Tool Deflection On the basis of the estimated cutting forces following Victor–Kienzle and also the actually measured forces, tool deflection was computed. Depending on the particular orientation, the values range between 1 and 5 lm and correspond well with the values found in the experiments [13].

5.2.3.5 Surface Quality The following observations on the milled surfaces of the hardended and annealed tool steel 1.2343 featuring a hardness of 52 HRC could be made (Fig. 5-17).

Fig. 5-16 Cutting forces at start of machining (sharp cutting edge) vs feed per cutting edge. Workpiece material 1.2343, 52 HRC and 42 HRC, vc = 30 m/min, ap = 10 lm

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Fig. 5-17 Hardened workpiece material (steel X38CrMoV5-1, 1.2343), surface roughness vs milling path, milling strategy (CL = climb milling, CO = conventional milling, Middle = middle of groove/50% ae) and feed per cutting edge (vc = 30 m/min)

The roughness of the milled surface of the hardened (52 HRC) tool steel ranges between Rz = 0.5 and 1.6 lm depending on feed per cutting edge and location of measurement. A distinct improvement of the surface roughness is, in contrast to experience when cutting with conventional cutting cross-sections, not noticeable during the decrease of the feed per cutting edge from 5 down to 0.5 lm. When milling the comparatively soft (42 HRC) workpiece material, surface roughnesses between Rz = 0.7 and 1.8 lm were found (Fig. 5-18). It can be seen that the topography formation of this material state after milling is much more stable during the progress of wear.

5.2 Micromilling

Fig. 5-18 Soft workpiece material (steel X38CrMoV5-1, 1.2343), surface roughness vs milled path, milling strategy (CL = climb milling, CO = conventional milling, Middle = middle of groove/50% ae) and feed per cutting edge (vc = 30 m/min)

5.2.3.6 Burr Formation Burr height ranges mostly from 5 to 60 lm. Only when milling with a feed per tooth of 0.5 lm do burr height values rise to the region of millimeters, whereas in the hardened material state (52 HRC) and at feed rates per cutting edge of ³3 lm almost a total absence of burrs can be achieved (Figs. 5-19 and 5-20). The influence of the milling strategy prevails: on the side of climb milling, higher burr formation can be observed. Furthermore, burr formation increases heavily with progressing wear. In addition, higher burr formation is found when milling the hard workpiece material (1.2343, 52 HRC) in comparison with the softer material (1.2343, 42 HRC) (Figs. 5-19 and 5-20), which can be attributed to the faster progress of wear and to the higher cutting forces related to the higher input of energy and leading to higher temperatures.

5.2.3.7 Process capability On the basis of an analysis method adapted from VDI-Directive 2851, statistical statements with regard to the accuracy were collected and a characterizing method that can be transferred to another milling machine for comparison purposes was developed. According to the minimum requirements for short-term capability analyses following VDMA Conformity Sheet 8669, the number of identical parts was fixed at 25. Table 5-7 gives information on the process parameters applied in this method.

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Fig. 5-19 Burr height vs progress of wear, milling strategy (CL = climb milling, CO = conventional milling) and feed per cutting edge, hardened material (1.2343, 52 HRC)

Fig. 5-20 Burr height vs progress of wear, milling strategy (CL = climb milling, CO = conventional milling) and feed per cutting edge, hardened material (1.2343, 42 HRC)

5.3 Conclusions Table 5-7 Test geometry and process parameters of the capability study

Parameter

Value

Feed per edge Cutting velocity Feed velocity Spindle speed

fz = 3 lm vc = 28 m/min vf = 180 mm/min n = 30 000 min–1

In order to achieve a capability index of cpk = 1.33, which is a standard value for machining processes, a dimensional accuracy of 0.02–0.05 mm has to be achieved. These tolerances could be readily maintained during the experiments, so that results obtained during the experiments can be regarded as well founded.

5.2.3.8 Manufacturing of Mold, Models and Test Structures for Demonstrator Components and Wear Analyses Knowledge gained from these research efforts was realized in the manufacture of workpieces which show a large variety of different structural details, e.g. a mold for a demonstrator planet wheel (Fig. 5-21).

5.3 Conclusions

It has been demonstrated that current micromilling tools still have severe problems to be overcome. Further miniaturization of these tools is limited by the process itself, the choice of the tool material and the ability to check the machined result. Another problem persists in the fact that coating has hardly been mastered by the manufacturers. Droplets and rather thick coatings prohibit machining with good surface roughness and marginal burr formation.

Fig. 5-21 Planet wheel with tooth flank in involute shape. X38CrMoV5-1 (1.2343), hardness 52 HRC

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So far, micromilling tools manufactured by EDM show the best results regarding tool life and surface topography. The tool life of EDM-manufactured tool was on average 2.5 times longer than that of the tools produced by laser ablation and 5 times longer than those created by ion beaming. Nevertheless, wear mechanisms severely changed the tool geometry. In further steps these tools will be miniaturized and coated to lessen tool wear. The choice of the matching workpiece material is another very important factor for machining results. Small grain sizes and a homogeneous distribution of second phases play a vital role in terms of tool life, process stability and machining result. This will be shown in Chapter 9.

5.4 References 1 R. Bierhals, K. Cuhls, V. Hüntrup,

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M. Schünemann, U. Thies, H. Weule, Mikrosystemtechnik – Wann kommt der Durchbruch?, Physica Verlag: Heidelberg, 1999. R. Bierhals, K. Cuhls, V. Hüntrup, M. Schünemann, H. Weule, U. Thies, Wirtschaftliche Potentiale der Miniaturisierung aus industrieller Sicht. Abschlussbericht zur Studie des Wirtschaftsministeriums Baden-Württemberg; Karlsruhe: Wirtschaftsministeriums Baden-Württemberg, 1999. J. Hesselbach, A. Raatz, J. Wrege, H. Herrmann, A. Wenda, H. Weule, C. Buchholz, H. Tritschler, M. Simon, F. Klocke, M. Weck, A. v. Klitzing, J. v. Bodenhausen, MikroPRO – Untersuchung zum internationalen Stand der Mikroproduktionstechnik. Schriftenreihe des Instituts für Werkzeugmaschinen und Fertigungstechnik der TU Braunschweig; Braunschweig: TU Braunschweig, 2002. G. Hollenstein, Formtechnik für Mikroteile und mikrostrukturierte Teile. In: Tagungsunterlagen IVAM-Workshop Industrielle Mikrosysteme aus Kunststoff, Dortmund, 2001. T. Masuzawa, State of the art of micromachining. Ann. CIRP 2000, 49, 1–16. T. Altan, B. Lilly, Manufacturing of dies and molds. Ann. CIRP 2001, 50, 1–24. H. K. Tönshoff, Technologieentwicklung im Werkzeugbau. In: Tagungsband Werkzeugbau-Kolloquium Hannover; 1997, pp. 105–130.

8 L. Alting, F. Kimura, H. N. Hansen,

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G. Bissacco, Micro engineering. Ann. CIRP, Vol 52, 2003. 1–23. J. H. Hernández-Camacho, Frästechnologie für Funktionsflächen im Formenbau; Dissertation, Universität Hannover, 1991. C. Ehrlich, Heißprägen für biomedizinische Anwendungen. Presented at MEDICA 2003, Mikrotechnik für die Medizintechnik, Düsseldorf; 2003. A. Menz, Mikrofluidische Komponenten aus Silizium, Anwendungen und Fertigungstechnologien. Presented at MEDICA 2003, Mikrotechnik für die Medizintechnik, Düsseldorf; 2003. H. Weule, J. Schmidt, V. Hüntrup, Micromilling of ferrous materials. Prod. Eng. 1999, 6 (2), 17–20. K. Weinert, C. Schwietering, Simulationsgestützte Mikrofräsbearbeitung – Steigerung der Prozesssicherheit bei der Herstellung mikrogefräster Formeinsätze. Werkstattstechnik – wt 2001, 91 (2), 76–79. J. Schmidt, H. Tritschler, Improvement of micro end milling tools through variation of tool manufacturing method and geometry. In: Proceedings of the 3rd Euspen International Conference, Eindhoven; 2002. N. Taniguchi, Current status in and future trends of ultraprecision machining and ultrafine materials processing. Ann. CIRP 1983, 32, 573–582. N. Ikawa, R. Donaldson, R. Komanduri, W. König, P. McKeon, T. Moriwaki, I. Stowers, Ultraprecision metal cutting

5.4 References

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18

19

20

21

22

23

24

25

26

27

– the past, the present and the future. Ann. CIRP 1991, 40, 587–594. K. Schubert, W. Bier, G. Linder, D. Seidel, Herstellung und Test von kompakten Mikrowärmeübertragern. Chem.-Ing.Tech. 1998, 61, 172–173. S. Asai, Y. Taguchi, K. Horio, T. Kasai, A. Kobayashi, Measuring the very small cutting-edge radius for a diamond tool using a new kind of SEM having two detectors, Ann. CIRP 1990, 39, 85–88. J. M. Castevens, Diamond turning of steel in carbon-saturated atmospheres, Precision Eng. 1983, 5, 9–11. C. Evans, J. B. Bryan. Cryogenic diamond turning of stainless steel. Ann. CIRP 1991, 40, 571—575. T. Moriwaki, Experimental analysis of ultraprecision machining. Int. J. Jpn. Soc. Precision Eng. 1995, 29, 287–290. T. Moriwaki, E. Shamoto, Ultraprecision diamond turning of stainless steel by applying ultrasonic vibration. Ann. CIRP 1991, 40, 559–562. E. Paul, J. C. Evans, et al., Chemical aspects of tool wear in single point diamond turning. Precision Eng. 1996, 18, 4–19. R. Rentsch, I. Inasaki, Investigation of Surface integrity by molecular dynamics simulation. Ann. CIRP 1995, 44. Z. J. Yuan, M. Zhou, et al., Effect of diamond tool sharpness on minimum cutting thickness and cutting surface integrity in ultraprecision machining. J. Mater. Process. Technol. 1996, 62, 327–330. T. Tanaka, N. Ikawa, et al., Affinity of diamond for metals. Ann. CIRP 1981, 30, 241–250. E. Brinksmeier, J. Dong, R. Gläbe, Diamond turning of steel molds for optical applications. In: Proceedings of the 4th

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31

32

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Euspen International Conference, Glasgow; 2004, pp. 155–157. G. Grathwohl, D. Godlinski, et al. Alumina tools for precision machining of steel. In: Proceedings of the 3rd Euspen International Conference, Eindhoven; 2002. J. Schmidt, H. Tritschler, Microcutting of steel. Microsyst. Technol. 2004, 10, 167– 174. P. Gumbsch, D. Löhe, J. Fleischer, J. Schmidt, V. Schulze, T. Hochrainer, L. Delonnoy, J. Kotschenreuther, Modeling, simulation and experimental tests for process scaling of cutting processes with geometrically defined edge. In: Proceedings of the 1st Colloquium of DFG Priority Program Process Scaling, Bremen; 2003. J. Schmidt, H. Tritschler, L. Bischoff, W. Schulz, Improvement of micro end milling tools through variation of tool manufacturing method and geometry. In: Proceedings of the ICMNS 2002, Kunming, China; 2002. J. Schmidt, D. Spath, J. Elsner, V. Hüntrup, H. Tritschler, Requirements of an industrially applicable microcutting process for steel microstructures, Microsyst. Technol. 2002, 8, 402–408. J. Fleischer, T. Masuzawa, J. Schmidt, M. Knoll, New applications for micro EDM. J. Mater. Process. Technol. 2004, 149, 246–249. J. Schmidt, H. Tritschler, H. Haberer, Cutting tools and material conditioning for micro end milling of tool steel. In: Proceedings of the 2nd Euspen International Conference, Turin; 2001. H. Victor, Beitrag zur Kenntnis der Schnittkräfte beim Drehen, Hobel und Bohren; Dissertation, Universität Hannover, 1956.

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6 3D Microstructuring of Mold Inserts by Laser-based Removal A. Stephen, F. Vollertsen, Bremen Institute of Applied Beam Technology (BIAS), University of Bremen, Germany Abstract The mass production of plastic microparts is mainly performed by low-cost replication processes such as injection molding (see Chapter 10) and hot embossing using precisely machined metallic molds. For microinjection molding of metals and ceramics and for injection molding of the plastic performs for microcasting (see Chapters 11, 12 and 13), metallic molds are also required. The structures contained in such molds can be either machined directly into the metal or the negative part into a polymer substrate which will subsequently be converted into a metallic mold by replication techniques. Lasers are a flexible tool for micromachining of both metals and polymers. The material removal can be achieved either by physical or chemical processes or by laser ablation or laser etching. In this chapter, an overview is given of the basic principles and technological requirements for the different laser removal technologies. Key aspects of the various machine concepts are reviewed, for example regarding laser sources and methods, and also materials aspects and processing strategies. Some examples are given of performance improvements with respect to processing time and surface quality. For the production of large-area molds including microstructures of different heights and shapes a process chain consisting of two laser processes, laser ablation of polymers and laser etching of metals, and an intermediate replication step of either injection molding or electroplating is presented. As an example of the process chain, the fabrication of a complexshaped metallic mold is described, which permits the production of microfluidic components optimized for hermetic sealing in special housings by ultrasonic welding.

Keywords laser ablation; laser etching; direct writing; mask projection

Advanced Micro and Nanosystems Vol. 3. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31208-0

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6.1 6.2 6.2.1 6.2.1.1 6.2.1.2 6.2.2 6.2.2.1 6.2.2.2 6.2.2.3 6.3 6.4 6.5

Introduction 132 Laser Removal Technologies 134 Laser Ablation 134 Focusing Technique 135 Projection Technique 140 Laser Etching 144 System Design 144 Process Fundamentals 145 Shaping Methods 149 Process Chains for Mold Fabrication Conclusion 156 References 158

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6.1 Introduction

The market for precision molded microparts is continuously growing and penetrating into new areas of application, e.g. microfluidic devices for biomedical analysis methods [1]. For the mass production of low-cost plastic, metallic or ceramic microparts by injection molding or by hot embossing, metallic inserts or forming tools with high surface quality and also often with structures of specific shapes are required to satisfy the operational performance of the molded component [2]. For many novel applications the molded parts contain microstructures of different size and shape. Often, positive (mesa) as well as negative (groove) structures are needed [3]. For molding of positive structures the corresponding negative counterpart has to be machined in the metallic mold, for example by processes leading to local material removal. Positive counter-structures in the mold can also be machined by this method; however, it is not efficient with regard to the processing time owing to the circumstance that the volume of the material to be removed is much larger than the volume of the structure itself. Such positive metallic structures can be efficiently fabricated by microstructuring of polymer substrates leading to a negative master which can subsequently be converted into a positive metallic form by replication, e.g. the LaserLIGA process first reported in 1995, and in a more applicable manner in 1996 [4, 5]. For the production of such manifold inserts, flexible manufacturing processes and technological methods are of interest. Performance, cost and environmental impact are important measures of processes for industrial applications. Material processing with lasers is an expanding field since it not only makes manufacturing cheaper, faster, cleaner and more accurate but also opens up entirely new technologies and manufacturing methods that are simply not available using standard techniques [6, 7]. In industry, lasers with nanosecond pulses, e.g. the Nd : YAG laser, are now recognized as a rapid and efficient tool for machining microfeatures into mate-

6.1 Introduction

rials that are hard to machine, such as metals and ceramics [8, 9]. The short pulse duration leads to high peak power intensities that remove most of the material via vaporization, resulting in an accuracy of a few micrometers with a layer-wise ablation and a very small heat-affected zone compared with longpulse laser material processing [10]. In order to produce microfeatures, the laser beam is focused to spot sizes of well below 50 lm. This requires a high-quality laser beam with an M2 value, which defines the deviation from an ideal Gaussian beam, not much larger than 1, i.e. a fundamental mode quality of the laser beam. A severe problem with laser ablation lies in the incompatibility of high ablation rates and good surface qualities. High ablation rates are usually connected with strong mechanical damage of the surface and significant heating of the material, leading also to serious melt formation due to the high laser power necessary. It could be shown that melting and, therefore, recast could be almost avoided with laser pulses in the femto- and picosecond range. However, the process speed is very slow compared with nanosecond applications owing to the low energies used [11]. Experimental investigations of machining metals have also shown that even femtosecond pulses can cause deleterious thermal effects [12]. Furthermore, amplified femtosecond laser systems are of limited use in industrial applications because they are complicated, less reliable and often not cost-effective. Their repetition rate is restricted to a few kilohertz and their average power to a few watts [13]. Therefore, femtosecond pulses cannot provide the best results in all cases and other longer pulse lasers may give a more complete solution in specific applications such as microstructured mold inserts [14]. Deleterious thermal effects, recast or redeposition of evaporated material can be greatly reduced by chemical-assisted laser machining in salt solutions [15] or even efficiently avoided by using laser activation of chemical etching reactions at the interface between the solid and a reactive fluid [6]. With the assistance of chemical reactions, a much lower surface temperature is required in order to activate the chemical material removal. Laser-induced wet chemical etching of metals can be achieved in different solutions of acids or bases using continuous-wave (cw) lasers in the visible or near-infrared region, e.g. the Nd : YAG laser [16]. The interaction of the focused laser beam with the material surface covered with a liquid etchant leads to an enhanced thermochemical etching reaction and thus to material removal from the irradiated area. In contrast to metals, precise and clean micromachining of polymers is possible by laser photoablation owing to the specific material properties such as low thermal conductivity and high photochemical sensitivity. For effective machining, high absorption of the material and short pulses and high repetition rates are necessary, which can be provided by lasers operating at wavelengths in the UV region, e.g. excimer lasers [17].

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6.2 Laser Removal Technologies

Laser machining refers to the controlled removal of material by laser-induced heating from the surface or bulk of the workpiece. In general, the domain for different laser material processing techniques as a function of laser power and interaction time can be divided into three major classes, namely involving only heating (without melting/vaporizing), melting (no vaporizing) and vaporizing [18]. Obviously, the laser power density and the interaction time are selected in each process so that the material concerned undergoes the desired degree of heating and phase transition. It is evident that laser-assisted chemical removal, which relies on surface heating without surface melting, requires low power density. On the other hand, surface glazing, which involves melting, requires a high power density. Similarly, ablation mainly removes material as vapor, and hence needs delivery of a substantially high power density within a very short interaction or pulse time. Solid-state lasers operated in the cw or pulsed mode, namely Nd : YAG lasers, cover all these domains, even if UV radiation is required as they can be frequency-tripled or -quadrupled. In general, shorter wavelengths lead to smaller spot sizes and to a better absorption efficiency. However, metals are less sensitive to laser wavelength and can be processed at all wavelengths. Nevertheless, excimer lasers are more suitable for processing if high pulse energies in the UV region are required [19]. An important challenge for all processes based on laser-based material removal is the control of the precision of the shape. This is because all methods are non-contact methods and the removal rate will be subject to certain influences depending on the exact processing conditions and material properties, making only limited prediction and direct control of the removal rate possible. In the following, both laser technologies, laser ablation and laser etching, will be described in more detail, particularly with regard to the fabrication of molding inserts. 6.2.1 Laser Ablation

The choice of the laser or the technique by which a particular application of laser ablation is accomplished is usually determined by the exact details of the application. In general, the most important parameters for choosing a laser microprocessing route are material, feature size, speed and resolution. The material largely determines the laser which can be used since most laser applications are fundamentally reliant on absorption of the laser light by the sample material. The required feature size, speed and resolution govern the choice of the processing technique and the systems design used. Bearing the above factors in mind, a processing method has to be chosen which can fall into two basic categories: focusing and projection techniques [7, 20]. Beam focusing is based on directly focusing the laser beam on the sample

6.2 Laser Removal Technologies

surface and using the focused light in combination with serial positioning for the structure generation on the sample. In contrast, projection techniques use the laser beam to illuminate a mask, which is then projected on to the sample surface to perform the microprocessing of one layer simultaneously. In general, serial machining is good for prototyping up to small series production because there is ablation of one pattern after another and data from a CAD file can (with some post-processing) be used to generate a tool path for the laser. On the other hand, mask projection is generally preferred for larger scale series production and the greater effort for the mask is economic for these high part numbers. These fundamentally different approaches are currently both used in industrial applications for the production of parts, but care has to be taken to utilize each method appropriately.

6.2.1.1 Focusing Technique Principle The beam focusing technique is performed as direct writing whereas the laser beam is focused to a small spot to define the feature dimension and then directed over the sample to machine the desired structure. Direct writing is widely used with many solid-state lasers, mostly the Nd : YAG laser. The fundamental reasons why this technique works well with these lasers are that they generally exhibit a very good beam profile and that the beam propagates in a well-defined Gaussian mode with low divergence. This enables the beam to be focused to a small, well-defined spot. One of the main advantages of direct writing systems is that they offer great flexibility in the structures that can be produced. The majority of direct writing systems maintain a fixed beam position with the sample being moved by using xyz-tables. Hence the features to be produced can simply be generated by the motion of the sample. This can easily be accomplished by CAD data input to the system to define the pattern to be machined. Hence the same system can be used to machine any structure provided that the sample can be moved appropriately. An extension of direct writing is the technique of beam scanning. In this case, the laser beam is moved rapidly over a small sample area by galvanometer scanning mirrors and the control of the scanning mirrors determines the features to be produced. The scanner mirrors can be controlled by CAD data files in the same way as the sample tables in direct writing. Beam scanning compared with direct writing using xy-tables shows much higher scanning speeds but poorer stability and accuracy. Direct writing or beam scanning, or combinations of both, are used widely in industry where cutting, drilling or profiling operations need to be performed. Rapid processing uses direct writing since the high repetition rate lasers permit very high machining speeds [21]. These speeds are still not fast enough for some areas, however, and direct writing can be combined with beam scanning to reduce the processing time further. This is achieved by processing an area

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covered by the scan field and then stepping the sample laterally by using xy-tables. Adjacent fields of the scanning mirrors have to be joined together to form the complete pattern. The ultra-high-speed scanner processing coupled with the large sample area provided by the xy-tables means that large sample numbers can be processed in a very short time.

System Design For micromachining application, a programmable galvanometer scanner guides the beam accurately with a lateral resolution of about 1 lm and at high speed over the target to create complex features, as mentioned above. A beam expansion telescope is integrated to increase the diameter of the beam in front of the scanhead before focusing for achieving smaller spot sizes and thus higher resolution. The set-up for this scanner system is shown in Fig. 6-1. In a pre-objective two-axis scanning arrangement, the laser beam passes through and is steered by a set of x and y mirrors that are coupled to galvanometers. The orthogonal arrangement of the x and y mirrors directs the beam down towards the workpiece and over the length and width of the scan field. Field distortion is compensated with an F-Theta lens after the two-mirror system. This permits both a large scan field (40 ´ 40 mm) and a small spot size of the laser beam at perpendicular incidence. The scanhead is mounted on a Z axis to adjust the focal plane with respect to the workpiece, which can be positioned by the XY stages in a field of 200 ´ 200 mm. The solid-state laser used in this set-up is a Q-switched Nd : YAG laser operating at a wavelength of 1064 nm in its fundamental Gaussian mode at a maximum output power of 16 W. Triggering permits pulsing operation up to 30 kHz. Since it delivers short pulses of about 100 ns, a high peak power of up

Fig. 6-1 Schematic view of a worksation for laser ablation by direct writing

6.2 Laser Removal Technologies

to 20 GW/cm2 can be achieved. The small M2 of < 1.2 allows the laser beam to be focused to spot sizes down to 18 lm in diameter for a focal length of 80 mm. Laser, scanner and stage interfaces are compatible for programming to permit automated part handling.

Structuring by Scanning The beam scanning over the workpiece generates a narrow trace of material removal. If the laser beam moves over the workpiece with a slight movement in lateral direction after the completion of each line, as shown in Fig. 6-2a for a bidirectional moving cycle, the whole surface of the workpiece or at least a certain part can be removed. To achieve uniform removal, the laser pulses and traces will be laterally overlapped in the range 50–80% with respect to the lateral dimension of the machined area by a single pulse. Nevertheless, this method leads to a residual waviness owing to the superposition of the traces. Among the laser parameters and the material, the roughness of the processed area significantly depends on the degree of overlap [22]. It is evident that the overlap of the pulses in each line is determined by the repetition rate and the scanning speed. To achieve the desired depth of the structure, the material will normally be removed layer by layer in several processing steps. Thereby, a residual waviness caused by scanning with fixed traces can almost be avoided by performing cross scans, as shown in Fig. 6-2 b, which will be rotated for each layer by a defined angle. The resulting pattern for five rotations of 158 each is shown in Fig. 6-2 c. According to this, the amplitude and the wavelength of the waviness can be reduced down to the size of the roughness. Furthermore, to compromise between high processing speeds and high surface qualities, the machining will be segmented into rough and finish machining. Rough machining will be performed using high laser powers to achieve high removal rates followed by applying low laser powers to finish machining with low roughness. Fig. 6-3 shows SEM images of surfaces after rough (a) and finish machining (b) of copper. For both processing steps equal repetition rates of 6 kHz and scanning speeds of 100 mm/s in addition to overlaps of 18 lm at a beam diameter of 18 lm are applied. The surface roughness measurements

Fig. 6-2 Scanning strategies: (a) bi-directional, (b) fixed cross and (c) wheel scan

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a)

b)

Fig. 6-3 Surface of a copper sample after (a) rough and (b) finish machining

were performed by using a stylus profilometer (Dektak 3ST) which determines the average roughness Ra and the mean peak to valley height Rz according to DIN 4768/1. The rough machining leads to a roughness Ra of 1.8 lm and an Rz of 10.4 lm using an average laser power of 7.5 W, which is reduced to a roughness Ra of 1.2 lm and an Rz of 6.9 lm by finish machining using a laser power of 4.6 W. By variation of the parameters such as scanning speed, overlap and average laser power for rough and finish machining, different values of roughness down to a few micrometers depending on the material can be realized. A process similar to surface hardening that leads to a further improvement in surface quality is performed whereby the beam intensity is increased to an extent such that not only is the transformation point for hardening reached but also melting takes place. In this case a small part of the material adjacent to its surface and the laser spot becomes molten and resolidifies after the laser beam has moved on. Owing to this melting and resolidification process, usually the crystal growth is suppressed due to rapid cooling and therefore a refined structure with improved quality, i.e. roughness, can be obtained. This process is referred to as ‘laser remelting’ or ‘laser polishing’ [23, 24]. For that purpose, the laser is operated in the cw mode. The machining result is determined for a specific material among the laser intensity and duration of exposure by transport phenomena inside the laser-induced melt pool at the surface of the workpiece. The surface of the copper sample in Fig. 6-3 after additional laser remelting or glazing is shown in Fig. 6-4. The processing leads to a surface roughness Ra of 1 lm and an Rz of 5.5 lm using a laser power of 9.8 W and a scanning speed of 0.1 mm/s. By applying a sequence of several laser finish machining and remelting steps, the roughness can be continuously decreased until a minimum value, which depends on the material, is asymptotically reached. A further decrease in rough-

6.2 Laser Removal Technologies Fig. 6-4 Surface of a copper sample after additional laser remelting

Table 6-1 Processing parameters for laser finishing/remelting of nickel surfaces

Step

Power (W) Repetition rate (kHz)

Overlap (lm)

Scan speed (mm/s)

Ra (lm)

Rz (lm)

Rough 1. Finish 1. Remelt 2. Finish 2. Remelt 3. Finish 3. Remelt 4. Finish 4. Remelt

8.0 1.2 9.8 1.2 9.8 1.2 9.8 1.2 9.8

10 5 10 5 10 2 10 2 10

100 150 0.3 150 0.3 60 0.3 60 0.3

4.0 2.3 1.7 1.9 1.3 1.6 1.1 1.4 1.0

21.7 12.6 9.2 9.9 6.5 9.0 5.7 7.7 5.6

10 30 cw 30 cw 30 cw 30 cw

ness is not possible by laser finish machining and/or remelting. The processing parameters of such a sequence for machining of electroplated nickel are given in Table 6-1. The corresponding values of roughness after each processing step are also given and in addition plotted in Fig. 6-5. As can be seen, the roughness achieved after each finish machining and remelting step depends on the initial roughness, i.e. the lower the initial roughness the better is the resulting roughness. A mold insert fabricated by laser ablation using this technique of direct writing is shown in Fig. 6-6. The letters of the writing ‘BIAS’ have different depth levels. The ‘B’ and ‘A’ have a height of 60 lm and the ‘I’ and ‘S’ a depth of 60 lm with respect to the ground level, which is 120 lm lower than the initial surface of the sample. Therefore, the maximum ablation depth of the structure is 180 lm. The measured ablation rate is 3 ´ 106 lm3/s for an applied pulse energy of 0.8 mJ equivalent to a pulse power of 6.5 kW used for rough machining.

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6 3D Microstructuring of Mold Inserts by Laser-based Removalpreading Fig. 6-5 Roughness on the lateral surfaces after several finish machining/remelting steps

Fig. 6-6 Light microscope image of a mold insert made of electroplated nickel produced by laser ablation and direct writing technique

6.2.1.2 Projection Technique Principle If the application requires the use of an excimer laser, e.g. large-area machining of polymers, then it may not be advantageous or even possible to use direct writing methods. The beam profile from standard excimer lasers does not lend itself easily to direct writing but the multi-mode nature of the excimer laser output does mean that it is a very good source for mask projection systems which require high pulse energies. Hence the majority of micromachining applications of excimer lasers use mask projection methods [25].

6.2 Laser Removal Technologies

In general, masks are made from chrome-on-quartz. The mask is projected on to the sample by a high-resolution lens system which also demagnifies the mask pattern. The mask can be mounted on xy-positioning tables and the sample can also be moved by in xyz-axes to align the sample and mask before the laser operation. The laser beam is kept fixed in position. The ability of an excimer laser to process a relatively large area in parallel is a consequence of the high pulse energy. A typical field size at the workpiece might be in the range of 1 ´ 1 mm, depending on the material and laser energy. Structures within this area are processed in parallel, allowing die-by-die exposure as in a lithographic stepper. Larger areas can also be processed with non-repeating patterns by synchronized motion of the mask and workpiece. One of the main advantages of mask projection is that the mask and workpiece can be moved independently of each other, which allows great flexibility in the range and types of features which can be micromachined. Mask projection systems can be used in the following ways to produce a variety of microstructures [26]: · Mask and workpiece stationary This is most commonly used for step-and-repeat processing where a single feature needs to be produced repeatedly several times on the sample surface. The beam area at the mask can illuminate many shapes at the same time and all of these can be machined at the same time and then the entire pattern is stepped and repeated. When a small aperture is used as a mask, this technique can also be applied for direct machining since this small projected circle acts as a focused laser beam, having the disadvantage of small efficiency between laser output and the laser energy used. · Moving mask This allows ramped and contoured structures to be produced by control of the mask position during laser firing. Applications include 3D micromachining. · Moving workpiece This can be used to produce slots or microchannels in the sample. A fixed shape at the mask defines the cross-section of the channel and the sample is moved to produce the structure with a defined path. Straight, curved or arbitrarily contoured channels can be produced. Application areas include microfluidic systems and biomedical devices. · Synchronized scanning Any large, non-repeating pattern can be transferred to the sample by synchronized scanning both the mask and sample under the laser beam, called a ‘dynamic projection system’. In this manner, high-resolution patterns can be produced over large areas. Applications include large-area microfluidic systems and biomedical devices. It should be noted that the same system can be used to produce all these features, highlighting the great flexibility of mask projection techniques. Typically both mask and workpiece are each mounted on precision motorized stages, so

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scanning operations can be performed automatically under computer numerical control. Other parameters can also be varied during the process, in particular the energy density and repetition rate of the laser.

System Design The basic concept with the key elements of a typical mask projection system is shown in Fig. 6-7. Before reaching the mask, the laser beam from a KrF excimer laser operating at a wavelength of 248 nm passes through a series of elements including a variable attenuator, a beam shaper and a homogenizer. The beam shaper simply transforms the beam cross-section, making it square to illuminate the homogenizer fully. The homogenizer then splits the beam into a large number of sources, each of which illuminates the mask from a different direction. The plane of uniformity of this beam is then made coincident with the plane in which the mask is positioned. This improves the uniformity of illumination at the mask plane, which is essential because of the poor quality of the raw laser beam, and also introduces off-axis components into the illumination. Off-axis illumination allows features with vertical and even undercut sidewalls to be produced. Such structures cannot be achieved with plane wave illumination. The illuminated mask is subsequently imaged by an objective lens on a projection scale of typically 5:1 on to the substrate. The mask and the polymeric substrate are synchronously moved perpendicular to the laser beam using two computer-controlled xy-translation stages, as shown schematically in Fig. 6-7. Some provision for alignment is included based on through-lens viewing using a CDD camera.

Vmask = –5 ·Vsubstrate

Fig. 6-7 Set-up for laser ablation by mask projection

6.2 Laser Removal Technologies

Structuring by Imaging An alternative to direct laser micromachining of metals is to adopt an approach where the laser is used to define a polymer master for subsequent replication in metal or if required even in plastic or ceramic. The polymer master is replicated, for instance, by an electroplating step in which the cavities are filled from the bottom with metal, usually nickel. By continuing the electroplating until a thick backing plate has grown over the entire surface, a tool for injection molding or hot embossing can be made. In principle, excimer lasers can machine a very wide range of materials [27]. However, they are particularly well suited to polymers as the typical wavelengths of excimer lasers provide photon energies up to 6.4 eV, which is sufficient for direct non-thermal cracking of polymer bonds leading to ablative photochemical decomposition (cold ablation). However, this process is characterized by a combination of photochemical and photothermal interactions. Hence, for most polymers strong absorption at excimer wavelengths ensures efficient laser–target coupling with low penetration depths of less than 1 lm, while the typical pulse duration of several nanoseconds and the relatively low thermal conductivity ensure minimal thermal diffusion during the ablation process and a very small heat-affected zone. In many cases, excellent surface finish and minimal damage (melting and/or debris) can be achieved. Most other materials cannot be machined with such quality by laser radiation. Metals, for example, typically exhibit very high ablation thresholds, as a consequence of high reflectivity and good thermal conductivity and usually cannot be machined without significant damage. The laser in combination with the demagnifying mask projection system allows the generation of complex miniaturized structures. The size of the micromachined area by excimer laser photoablation in conventional static projection systems is limited owing to the available laser power and the maximum size which can be imaged by the objective lens with high resolution. With the dynamic projection system, which allows synchronized scanning, sequential machining of small parts of the structure leads to a large area of microstructures of several cm2 limited only by the travel range of the stages. The synchronous scanning of the mask and the substrate in the opposite direction and perpendicular to the laser beam results in a lateral resolution of < 3 lm. Microstructuring of polymers is performed in this set-up at repetition rates of 50 Hz and an energy density of 1 J/cm2 at the sample surface. To reduce redeposition of ablated material, the sample is flushed with compressed air and the particles are removed by an extraction system while processing. Figs. 6-8 and 6-9 show a part and a detailed view of a 10 ´ 20 mm sized microstructure in pure polycarbonate achieved by using the technology of synchronous scanning. The depth of the structure is 60 lm and the process depending angle of the walls 158. The machining by photoablation results in very smooth surfaces with a roughness Rz of about 50 nm. However, a residual waviness of about 0.5 lm due to the scanning process can be observed. The ablation rate is 0.25 lm per pulse, leading to a removal rate of 107 lm3/s.

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Fig. 6-8 Part of a large-area microstructure in polycarbonate

Fig. 6-9 Detailed view of Fig. 6-8

6.2.2 Laser Etching

Laser etching is based on a local thermal activation of chemical etching reactions on the surface of the material. In the following, the process fundamentals based on the system design and some aspects of shaping methods for the machining of molding inserts are described. 6.2.2.1 System Design The beam of a cw Nd : YAG laser operating at a wavelength of 1064 nm in its fundamental Gaussian mode at a maximum power of 16 W is focused to an estimated focal spot diameter of about 20 lm on the metal surface. Metals used are nickel and steel or for basic investigations also nickel alloys, which show the same processing characteristics. The workpiece is immersed in a liquid etchant consisting of 5.6 M phosphoric acid (H3PO4) and 1.5 M sulfuric acid (H2SO4). To perform laser-induced liquid-phase jet-chemical etching, a special liquidphase etching cell is integrated as shown schematically in Fig. 6-10. The cell consists of two parts, a co-axial nozzle assembly and a basin. The nozzle can be adjusted laterally and in height with respect to the laser beam. The etch liquid enters the nozzle tip in such a way that a swirl is given the liquid flow and is injected co-axially to the laser beam directly into the irradiated area. Four different nozzle designs with diameters of the tip ranging from 0.5 to 2 mm lead to flow-rates of the etchant between 2 and 20 m/s. The basin holds the workpiece and is mounted on a computer-controlled xyz-stage allowing a relative movement over an area of 100 ´ 100 mm at a resolution of 0.1 lm to position the workpiece with respect to the laser beam. Additionally, a usual three-electrode configuration is integrated into the chemical cell to perform electrochemical enhancement of the reaction if required.

6.2 Laser Removal Technologies

Fig. 6-10 Principle scheme of laser etching

The working electrode is the metallic workpiece itself connected by an insulated wire to the potentiometer. The counter-electrode is a circular formed wire made of inert platinum and is located upside of the workpiece in front of the nozzle. The reference electrode is a saturated calomel electrode (SCE) and located inside the basin.

6.2.2.2 Process Fundamentals Principle At room temperature many metals are protected against corrosion by a thin natural oxide layer on the surface and in many aggressive media behave like a noble metal. So, for example, in phosphoric acid negligible corrosion rates of < 10–8 lm/s at room temperature for titanium can be observed [28]. An increase in temperature results in a shift of the chemical equilibrium towards the formation of soluble metal salts and hydrogen. Time-resolved measurements of the electrical potential against an electrochemical reference electrode show that localized heating of the passivated metal by focused laser radiation results, in analogy with thermal corrosion, in localized dissolution of the passivation layer followed by chemical etching of the metal indicated by a sudden drop of the voltage upon laser illumination as shown in Fig. 6-11. After the end of laser irradiation a sudden increase in the electrical potential reveals an immediate interruption of the etching reaction due to repassivation of the metal surface which is essential for high-quality machining. In particular, on laser irradiation, temperatures much higher than the boilingpoint of the liquid can be reached on the surface in the zone of laser action. At such high temperatures, etch rates several orders of magnitude higher are measured than at the boiling-point of the etchant. The measured exponential Arrhe-

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6 3D Microstructuring of Mold Inserts by Laser-based Removalpreading Fig. 6-11 Temporal evolution of the potential for an NiTi alloy switching the later on/off

nius-type dependence of the laser-induced etch rate on laser power also supports the thermal nature of the chemical process [28].

Laser-induced Jet-chemical Etching In contrast to processing in gaseous media, liquid-phase chemical etching benefits from the high density of reactants in the liquid. Since etching processes are transport limited and an increase in the metal load of the etchant leads to a decrease in the etch rate, a fast exchange of the reaction products with fresh reactants is an essential requirement for avoiding saturation effects of the etch rate. Such efficient exchange is realized by the liquid jet-stream. It leads to an improvement of the processing speed and treatment quality due to the intensive mass transport and cooling of the workpiece by direct injection of the etchant into the laser-irradiated area. Fig. 6-12 a represents the dependence of the etch rate on the etchant flow-rate and laser power for a nickel–titanium alloy. It shows that etch rates up to 8 ´ 104lm3/s can be achieved for laser powers up to

flow-rate (m/s)

a)

b)

Fig. 6-12 (a) Dependence of etch rate on flow-rate and laser power. (b) Microphotographs of some corresponding grooves

6.2 Laser Removal Technologies

7 W and a flow-rate of 2 m/s. Compared with this, the etch rate at a flow-rate of 20 m/s is only half as much. This is due to the higher cooling effect by the liquid jet-stream, thus leading to lower thermal activation of the metal and hence to weaker etching reactions. At the same time, an enhancement of the etchant flow-rate improves the shape fidelity and treatment quality. The microphotographs of machined grooves at three characteristic processing conditions shown in Fig. 6-12 b illustrate this effect: No. 3 shows the same high quality as No. 1 in terms of a shape fidelity of 3 lm at a much higher etch rate which is equal to No. 2 showing a worse quality with a shape fidelity of only 12 lm. Thus, an increase in processing speed at equal quality can be achieved by simultaneously increasing the laser power and flow-rate.

Laser-induced Jet-electrochemical Etching Owing to the ionic nature of the reactants, the chemical etching reaction can be additionally influenced by an external electric field, applied between the metal workpiece and an additional electrode. In contrast to laser-assisted electrochemical machining, the applied voltages are very low (a few volts compared with more than 100 V) and the material removal, therefore, is not electrochemically dominated [29]. Laser-induced electric currents can be observed for a nickel–titanium alloy both in the cathodic, active and in the passivation region, as shown in Fig. 6-13. The results were obtained by a cathodic scan of the I/U characteristics and repeatedly switching the laser on and off. In this case, laser-induced currents are due to electrochemical etching reactions caused by thermal activation of the anodic dissolution and/or laser-induced breakthrough of the passivation layer. The corresponding etch rates are shown in Fig. 6-14 a using a laser power of 5 W and a flow-rate of 10 m/s. Depending on the polarity of the applied electric field, a significant increase in the etch rate (in the cathodic region) or improvement of the treatment quality (in the passivation region) can be observed as shown in Fig. 6-14 b. It is assumed that this is caused by an electro-

Fig. 6-13 Laser-induced electrical currents

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6 3D Microstructuring of Mold Inserts by Laser-based Removalpreading

a)

b)

Fig. 6-14 (a) Dependence of etch rate on applied voltage. (b) Microphotographs of some corresponding etched holes

Fig. 6-15 I/U characteristic of tool steel 1.3343 in sulfuric acid

chemical reduction of the passive layer towards cathodic potentials enhancing the chemical dissolution. In contrast to nickel or its alloys, hardened tool steel, e.g. 1.3343, is not protected by a thin native oxide layer at room temperature and rapidly corrodes in phosphoric and/or sulfuric acid. Nevertheless, it can be electrochemically passivated and, therefore, protected against corrosion outside the laser-irradiated area by applying an external electric field. The necessary voltage for achieving passivation is determined by scanning the current–voltage characteristic of the system as shown in Fig. 6-15 for sulfuric acid. Starting with a cathodic scan, the curve passes several active regions until the transpassive condition is reached at about 2 V. By scanning back, a broad passivation region can be observed until at 0.3 V an abrupt increase in the current indicates an instant activation of the metal dissolution and corrosion. When the tool steel is activated once, the characteristic hysteresis curve has to be scanned

6.2 Laser Removal Technologies

again for repassivation. In the passivation region, only low currents exist which decrease with decreasing voltage. Taking this and possible slight local and/or temporal fluctuations of the potential into account, the optimal processing voltage is about 0.5 V. Thus, combining the influence of three independent process parameters (laser power, etchant flow-rate and electric field), optimum conditions for obtaining predetermined treatment results can be found.

6.2.2.3 Shaping Methods For the laser etching, the achievable structures are not limited in maximum size because they are generated successively which, however, results in long processing times. An advantage of this technology is the possibility of generating structures with defined shape. Grooves can be machined by moving the workpiece perpendicular to the laser beam. Owing to the thermal activation of chemical reactions, the width and depth of the grooves are determined by the temperature distribution on the surface and the duration of the temperature rise [30]. Therefore, the shape of the groove reflects the intensity distribution of the incident laser beam and the effective heat conductivity of the machined material. In Fig. 6-16, the dependence of the depth of etched grooves in a 200 lm thick nickel–titanium foil on the energy per unit length obtained by varying the laser power at different scanning speeds is represented. In the parameter range investigated, an approximately linear dependence of etched depth on energy per unit length is observed. Therefore, the same depth can be achieved at even higher processing speeds by increased laser power and scanning speed. A groove which is machined using a laser power of 3 W and a moving speed of 10 lm/s is shown in Fig. 6-17. The depth and width achieved are approximately equal corresponding to an aspect ratio of 1. Because of the lateral heat diffusion from the zone of laser action, high aspect ratios cannot be realized by single scanning of the groove. The width and depth of the grooves simultaneously increase with increasing laser power. However, higher aspect ratios can be realized by multiple scanning of the laser beam along the same groove. The groove

Fig. 6-16 Dependence of etch depth on energy per unit length

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6 3D Microstructuring of Mold Inserts by Laser-based Removalpreading Fig. 6-17 Microphotograph of a machined groove

Fig. 6-18 Predetermined overlaps for a V-shaped groove

width is almost independent of the number of scans as the temperature increase is confined to the bottom of the groove, making it a quasi-anisotropic process. This leads to a continuously increasing depth and, in consequence, to higher aspect ratios (etched depth to width) of up to 10 [30]. To machine shape-controlled structures, the method for achieving high aspect ratios by multiple scanning of the laser beam along the same groove is expanded by a simultaneous lateral shift of every scan with respect to the groove axis. By controlling the overlaps between the laser-irradiated areas, the etched depth and hence the structure‘s shape can be determined [31]. For instance, Fig. 6-18 shows the calculated overlaps for a V-shaped groove with defined side angle as a function of the distance from its center. Using these parameters, a 150 lm deep and 200 lm broad V-shaped groove with a radius of the tip of about 5 lm is fabricated at a scanning speed of 10 lm/s and a laser power of 8 W. Fig. 6-19 shows an SEM image of the groove.

6.2 Laser Removal Technologies Fig. 6-19 V-shaped groove machined by multiple scanning in electroplated nickel

Fig. 6-20 Cross-section of a groove in electrolated nickel machined by projection technology

The measured surface roughness Rz inside the laser-irradiated area perpendicular to the groove axis is < 1 lm. Furthermore, a defined shape of the structures can be achieved by forming the intensity distribution of the laser beam on the sample surface by a projection system using an aperture of special shape. The main advantage of this method is that the microstructure can be machined with a single scan using high-power lasers resulting in high processing speeds. For example, Fig. 6-20 shows the cross-section of a groove achieved by projecting a right-angled triangle aperture on to the sample‘s surface with a laser power of 140 W at a moving speed of 10 lm/s. However, the resolution of such machined structures is much lower than for multiple scanning. Specifically, the minimum radius that can be achieved by projection technology using a high-

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6 3D Microstructuring of Mold Inserts by Laser-based Removalpreading

power laser is in the range of 25 lm compared with 5 lm for multiple scanning using a focused Gaussian laser beam of 20 lm in diameter.

6.3 Process Chains for Mold Fabrication

Shaping of metallic components for mold fabrication or rapid prototyping by laser processes such as laser-based removal or laser forming is an effective and efficient manufacturing technology as these methods require no mechanical contact and hence offer many of the advantages of flexibility in process chains [32]. The principle of an exemplary process chain is shown schematically in Fig. 6-21. It consists of two laser processing steps in combination with an intermediate step of micrometal injection molding (lMIM) or electroplating [33, 34]. Three steps lead to the metallic mold which can be used in a fourth step to fabricate plastic microparts [35]: 1. A pure polycarbonate substrate is microstructured by excimer laser ablation in a dynamic projection system (flying mask). This step offers the possibility of fast and low-cost fabrication of large-area master structures. 2. By lMIM of metals such as iron or electroforming of nickel, after depositing a starting layer, the polymeric master is converted into a metallic form. Subsequently the polymer can be removed mechanically or by chemical dissolution. 3. Additional structures are micromachined by laser etching. These structures show different shapes and heights compared with the microstructures of the

Fig. 6-21 Principle of an exemplary laser process chain

6.3 Process Chains for Mold Fabrication

master, for instance V-shaped grooves, and are easier to be machined directly into the metallic insert. Furthermore, the overall height of the microstructures can be increased to several hundred micrometers, which is difficult to realize by conventional techniques such as lithographic processes. 4. The finished metallic mold can be used for a fabrication of several hundred thousand similar plastic microparts by injection molding or hot embossing, resulting in low processing costs for each part. The two different laser processes combined by the intermediate processing step of replication lead to complex-shaped microstructures in metallic inserts. By replication in metal such as by lMIM or electroplating, the polymeric semi-master machined by excimer laser ablation is converted into a metallic mold containing the required positive structures. Subsequently, additional negative structures are machined by laser etching. Fig. 6-22 shows a master in polycarbonate which is designed for hot embossing of low-cost microfluidic components in polymers after replication in iron by lMIM as shown in Fig. 6-23. The replication by lMIM is performed by injecting an extremely fine metal powder mixed with an organic binder into the insert, removing the binder and sintering the powder to full density incorporating a certain shrinkage of about 15%. Before sintering, the polymer substrate can be removed and used again. The main influencing parameter on the quality of reproduction with respect to surface roughness and aspect ratio is, apart from the quality of the mold itself, the powder used. Generally, a smaller particle size of the powder allows higher aspect ratios and smoother surfaces. The reproduction in iron leads to a maximum aspect ratio of about 5 : 1 and a roughness Ra of 0.3 lm and an Rz of 2 lm using a powder of 2 lm particle size [33]. An alternative approach for replication is electroplating in nickel. As the excimer laser ablation is performed in pure polymer, a metallic starting layer first

Fig. 6-22 Master in polycarbonate produced by excimer laser ablation

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6 3D Microstructuring of Mold Inserts by Laser-based Removalpreading Fig. 6-23 Reproduction in iron by lMIM of the master

Fig. 6-24 Part of a large-area semimaster in pure polycarbonate produced by excimer laser ablation

has to be deposited on the sample. Fig. 6-24 shows the large-area master structure with an overall area of 2 cm2 in polycarbonate plated with a gold layer. It is designed for microfluidic applications hermetically sealed in housings. Fig. 6-25 shows the insert produced by electroforming of the master structure, after chemical dissolution of the polymer substrate. Therefore, the polymer master can be used only once. The initial roughness Rz of about 50 nm and residual waviness of about 0.5 lm due to the excimer laser ablation and scanning process are reproduced in the replication. The maximum aspect ratio which can be achieved by electroplating for these metal-coated substrates is limited to 5 : 1 as the isotropically conductive substrate shows an undirected growth behavior. As a consequence, the galvanic deposit will overgrow the apertures of contained cavities before they are completely filled with metal and partially filled cavities are left in the product. To achieve higher aspect ratios, even if this is not re-

6.3 Process Chains for Mold Fabrication Fig. 6-25 Electroplated insert of nickel with additional structures machined by laser etching

Fig. 6-26 Molded part in plastics fabricated by hot embossing of the electroplated insert

quired for the application shown, methods of two-component preforms can be applied to start metal deposition on the substrate bottom, i.e. cavities, only [34]. Additional machining by thermochemical etching leads to V-shaped grooves which form energy directors for ultrasonic welding by molding in plastics. They are used to hermetically seal the housing by a cover. Typically, the height of the grooves varies between 50 and 200 lm, the dihedral angle between 60 and 1208 and the length can be more than 100 mm with tolerances of a few micrometers. Fig. 6-25 shows two 400 lm wide and 200 lm deep V-shaped grooves of 20 mm in length. They are located on both sides close to the microfluidic structures. The roughness Rz perpendicular to the grooves is < 1 lm. Fig. 6-26 shows a molded part in plastics fabricated by hot embossing of the metallic insert. The laser processing time for the combined method in comparison with machining exclusively by excimer laser removal or laser-induced etch-

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6 3D Microstructuring of Mold Inserts by Laser-based Removalpreading Fig. 6-27 Laser processing times for the structures

single processes

Fig. 6-28 Cross-secton of a hermetically sealed microfluidic device by ultrasonic welding

ing is shown in Fig. 6-27. It clearly shows the large time savings made when using the process chain instead of a single process. An additional advantage of the process chain is the possibility of manufacturing complex-shaped structures as demonstrated. In a final processing step, this microfluidic device is joined with a plastic plate on top of it in order to hermetically seal the microfluidic structures. This is done by ultrasonic welding. Thereby, the ultrasonic energy is concentrated at the sharp cone end of the energy director, locally resulting in high temperatures and, in consequence, melting of the plastics. Fig. 6-28 shows the cross-section of such a joint.

6.4 Conclusion

High-resolution laser micromachining techniques which are suitable for the production of molds have been described. These techniques can be mainly divided into processes based on physical or chemical laser removal, and pulsed la-

6.4 Conclusion

157

ser ablation and continuous-wave laser etching. Well-established lasers for these applications are excimer lasers and solid-state lasers such as the Nd : YAG laser. For micromachining, both lasers show good results for feature sizes in the 5– 100 lm range. However, the best choice of laser depends on the material being machined and the processing strategy applied. In particular, detailed descriptions of Nd : YAG laser direct writing and excimer laser mask projection systems have been presented. Pulsed Nd : YAG laser ablation approaches higher processing speeds by utilizing the benefits of solid-state laser technology such as short pulses and high repetition rate operation. The excellent beam quality of the lasers leads to extremely high peak power densities that are a precondition for the machining of hard materials such as metals and allows precise machining with aspect ratios up to 50 : 1. If a UV wavelength is required, only excimer lasers deliver high UV power and can economically reach short wavelengths such as 248 or 193 nm. The beam characteristics make the excimer the ideal laser source for large-area machining with high resolution at moderate energy density as required for polymers. As an application this technology is used to manufacture microfluidic master structures for molding inserts which will be produced by replication techniques such as metal injection molding or electroforming of the master leading to a metallic copy. The machined area is several cm2 at a lateral resolution better than 3 lm. The remaining waviness due to the scanning process is < 0.5 lm. In contrast to laser ablation, laser chemical machining benefits from the low processing temperatures and debris-free material removal due to laser-induced chemical etching reactions. The achieved sidewall angle is about 38 and the surface roughness is < 0.3 lm. Furthermore, two processing strategies for the purpose of machining with defined structure shape are demonstrated by means of a V-shaped groove. Grooves with depths up to 200 lm and dihedral angles of 908 are realized. The determined roughness Rz is < 1 lm.

Table 6-2 Comparison of the different laser processes in the field of micromolding

Process

Laser

Material

Laser ablation focusing Laser ablation projection Laser etching

Q-switched Nd : YAG Excimer

Any me- 100 mm3/ 50 : 1 tals min Polymers 10–1 mm3/ 5 : 1 min Metals 10–2 mm3/ 10 : 1 min (nickel, steel)

cw Nd : YAG

Removal rate

Aspect ratio

Min. feature size

Height Roughness Thermal tolerance Ra influence

50 lm

5 lm

1 lm

High

3 lm

0.3 lm

20 lm

Low

10 lm

3 lm

0.3 lm

None

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In Table 6-2, a comparison of laser ablation of metals using the focusing technique, laser ablation of polymers using the projection technique and laser etching of metals with respect to processing time and quality aspects is given. As an application in molding technology, a process chain has been presented for manufacturing precise metallic molds by laser processing. First, excimer laser photoablation of a polymer substrate is performed in a dynamic projection system to achieve large-area master structures for metal injection molding or electroplating. Second, the as-produced metallic insert is additionally micromachined by laser-induced jet-chemical etching leading to complex-shaped microstructures with high accuracy. The process chain permits a significant decrease in processing time. The combined processing route is used to fabricate a mold allowing the production of microfluidic devices in special housings which can be hermetically sealed by ultrasonic welding.

6.5 References 1 T. Hanemann, W. Pfleging, J. Hausselt,

2

3

4

5

6 7

K.-H. Zum Gahr, Laser micromaching and light induced reaction injection molding as suitable process sequence for the rapid fabrication of microcomponents. Microsyst. Technol. 2002, 7, 209– 214. T. Hanemann, M. Heckele, V. Piotter, Current status of micromolding technology. Polym. News 2000, 25, 224–229. A. Stephen, T. Lilienkamp, S. Metev, G. Sepold, C. Schön, R.-P. Peters, Fabrication of large-area precision molds by laser-assisted micromachining. In: Lasers in Manufacturing, German Scientific Laser Society (ed.); Stuttgart: AT-Fachverlag, 2001, pp. 342–350. J. Arnold, U. Dasbach, W. Ehrfeld, K. Hesch, H. Löwe, Combination of excimer laser micromachining and replication processes suited for large scale production. Appl. Surf. Sci. 1995, 86, 251– 258. R. A. Lawes, A. S. Holmes, F. N. Goodall, The formation of moulds for 3D microstructures using excimer laser ablation. Microsyst. Technol. 1996, 3, 17–19. D. Bäuerle, Laser Processing and Chemistry; Berlin: Springer, 2000. S. Metev, V. Veiko, Laser-assisted Microtechnology; Berlin: Springer, 1998.

8 L. Herbst, J. P. Quitter, G. M. Ray, T.

9

10

11

12

13

14

15

Kuntze, A. O. Wiessner, S. V. Govorkov, M. Heglin, High peak power solid state laser for micromachining of hard materials. Proc. SPIE 2003, 4968, 134–142. A. Ostendorf, K. Körber, T. Nether, T. Temme, Material processing applications for diode pumped solid state lasers. Lambda Highlights 2000, 57, 1–2. M. Schmidt, G. Esser, Laser microprocessing – facts and trends. Proc. SPIE 2003, 5063, 163–171. M. Weikert, C. Föhl, F. Dausinger, T. Abeln, Surface structuring of metals with short and ultrashort laser pulses. Proc. SPIE 2003, 5063, 208–213. F. Dausinger, Femtosecond technology for precision manufacturing: fundamental and technical aspects. RIKEN Rev. 2003, 50, 77–82. T. Herrmann, B. Klimt, F. Siegel, Slashing the costs of high-precision micromachining. Photonics Spectra 2004, 38 (6), 70–72. N. H. Rizvi, Femtosecond laser micromachining: current status and applications. RIKEN Rev. 2003, 50, 107–112. L. Li, C. Achara, Chemical assisted laser machining for the minimization of recast and heat affected zone. Ann. CIRP 2004, 53 (1), 175–178.

6.5 References 16 R. J. von Gutfeld, E. E. Tynan, R. L. Mel-

17

18 19

20

21

22

23

24

25

26

27

cher, S. E. Blum, Laser enhanced electroplating and maskless pattern generation. Appl. Phys. Lett. 1979, 35, 651–653. T. Lippert, M. Hauer, C. R. Phipps, A. Wokaun, Fundamentals and applications of polymers designed for laser ablation. Appl. Phys. A 2003, 77, 259–264. W. M. Steen, Laser Material Processing; New York: Springer, 1991. R. Paetzel, Comparison excimer laser – solid state laser. In: Proceedings of ICALEO, E. Beyer (ed.); Orlando, FL: LIA, Laser Inst. of America, 2002, CD-ROM. M. C. Gower, N. H. Rizvi, Applications of laser ablation to microengineering. Proc. SPIE 2000, 4065, 452–460. T. Hanemann, J. Hausselt, R. Ruprecht, M. Skrifvars, W. Pfleging, K.-H. Zum Gahr, Rapid fabrication of microcomponents. Proc. SPIE 2000, 4019, 436–443. J. Koch, S. Metev, G. Sepold, Lasergestützte Hochgeschwindigkeitsbearbeitung für die Kleinserienfertigung von 3D-Mikrobauteilen. DFG Abschlusskolloquium SPP 1012; 2003. D. Schuöcker, High Power Lasers in Production Engineering; Singapore: World Scientific, 1999. R. Poprawe, W. Schulz, Development and application of new high-power laser beam sources. RIKEN Rev. 2003, 50, 3–10. M. C. Gower, Excimer lasers: principles of operation and equipment. In: Laser Processing in Manufacturing, R. C. Crafer, P. J. Oakley (eds.); London: Chapman and Hall, 1994, pp. 189. N. H. Rizvi, Production of novel 3D microstructures using excimer laser mask projection techniques. Proc. SPIE 1999, 3680, 546–552. M. C. Gower, Excimer lasers: current and future applications in industry and

28

29

30

31

32

33

34

35

medicine. In: Laser Processing in Manufacturing, R. C. Crafer, P. J. Oakley (eds.); London: Chapman and Hall, 1994, 262– 272 R. Nowak, S. Metev, Thermochemical laser etching of stainless steel and titanium in liquids. Appl. Phys. A 1996, 63, 133–138. A. K. M. De Silva, P. T. Pajak, D. K. Harrison, J. A. McGeough, Modelling and experimental investigation of laser assisted jet electrochemical machining. Ann. CIRP 2004, 53 (1), 179–182. A. Stephen, T. Lilienkamp, S. Metev, G. Sepold, Laser-assisted chemical micromachining of metals and alloys. RIKEN Rev. 2002, 43, 56–62. A. Stephen, G. Sepold, S. Metev, F. Vollertsen, Laser-induced liquid-phase jetchemical etching of metals. J. Mater. Process. Technol. 2004, 149, 536–540. F. Vollertsen, Laserstrahlumformen – Lasergestützte Formgebung: Verfahren, Mechanismen, Modellierung; Bamberg: Meisenbach, 1996. A. Rota, T. V. Duong, T. Hartwig, Micropowder metallurgy for the replicative production of metallic microstructures. Microsyst. Technol. 2002, 8, 323–325. V. Piotter, N. Holstein, K. Plewa, R. Pupcrecht, J. Hausselt, Replication of microcomponents by different variants of injection molding. Microsyst. Technol. 2004, 10, 547–551. F. Vollertsen, A. Stephen, S. Metev, A process chain for laser-assisted manufacturing of precision metallic moulds. In: Micro Engineering, H. N. Hansen, L. Alting (eds.); Copenhagen, Denmark: IPL TU, 2003, pp. 49–52.

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7 Micro-EDM for Mold Inserts J. Schmidt, M. Knoll, Institute of Production Science (wbk), University of Karlsruhe (TH), Germany T. Masuzawa, Institute of Industrial Science (IIS), University of Tokyo, Japan Abstract Micro-EDM (electrodischarge machining) is a promising technology for micromachining. In combination with milling and laser ablation, a large variety of structures can be produced. In this chapter, potential applications of microEDM for the production of mold inserts are presented. The fundamentals of the process and special variants of EDM such as die-sinking EDM, wire EDM, WEDG (wire electrodischarge grinding) and EDM milling are described. The importance of this technology for the Collaborative Research Center 499 regarding mold insert production is explained, referring to die-sinking EDM and EDM milling. EDM milling is difficult to use because it demands compensation of the wear of conventional electrodes. A new type of electrode was developed at the Institute of Production Science (wbk) to overcome this difficulty. Finally, new application fields for micro-EDM are identified.

Keywords electrodischarge machining; micro-EDM; EDM milling; WEDG; wire electrodischarge grinding; micromolds

7.1 7.1.1 7.1.2 7.1.3 7.1.4

Introduction

162

7.1.5 7.1.6 7.1.7

Fundamentals of EDM 163 Micro-EDM 164 Die-Sinking EDM 165 Wire EDM 165 WEDG 166 EDM Milling 167 Synopsis of Micro-EDM Technologies 170

7.2 7.3 7.4

State-of-the-Art in the Industrial Environment 170 Mold Insert Production by Die-Sinking EDM 172 Mold Insert Production by EDM Milling 173

Advanced Micro and Nanosystems Vol. 3. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31208-0

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7.5 7.6 7.6.1 7.6.2 7.6.3 7.7 7.8

Development of a New Type of Electrode for EDM Milling Special WEDG Applications for Micro-EDM 176 EDM of Milling Tools 179 EDM Lathe for Micromachining 179 Micropunching with WEDG Support 180 Conclusion 182 References 183

174

7.1 Introduction

Replication techniques (see Part 3) are of the utmost importance for the economical production of microparts. Major challenges for these technologies are on the one hand the optimization of the process parameters and on the other hand the production of the necessary tools and dies. The use of machining technologies from mechanical engineering such as cutting, milling and grinding allows for the machining of wear-resistant materials at acceptable cost. Today there is still little experience concerning the application limits of these technologies. In all material removal processes, the requirements for microparts are small unit removal and high precision. Compared with processes such as cutting and grinding, EDM is one of the most important methods concerning the variety of machinable shapes and machining accuracy. The advantages of EDM are due to its basic characteristics: · All surfaces of the tool electrode can be used for material removal, whereas only the cutting edges are used in the cutting process. · EDM is an almost force-free process whereas cutting force is necessary for material removal in cutting processes. EDM can be divided into the following applications: · Die-sinking EDM (SEDM). · Wire EDM (WEDM). · (Wire) electrodischargegrinding (WEDG, EDG). · EDM milling. Comparatively few institutions are carrying out research in the field of EDM, and are mostly located in Japan, Germany and Switzerland. In Germany only a small number of institutes are conducting research in the field of EDM, for example, the Laboratory for Machine Tools and Production Engineering (WZL, Aachen), the Institute for Machine Tools and Factory Management (IWF, Berlin), the Institute of Production Technology (IPT, Dresden), the Institute for Fundamental Electrical Engineering and Electromagnetic Compatibility (IGET, Magdeburg) and the Institute of Production Science (wbk, Karlsruhe). The research topics vary from the development and optimization of machine equipment (e.g. generators) to the optimization of the process itself.

7.1 Introduction

7.1.1 Fundamentals of EDM

The principle of EDM is simple. A voltage is impressed on two electrodes (one tool electrode and one workpiece electrode), which are separated from each other by a dielectric fluid (Fig. 7-1). The two electrodes are moved close together and if the voltage is high enough (> 80 V), the dielectric fluid will break down and conduct an electrical current, causing an electrical discharge (spark) between them. The sparks will produce an extremely high temperature (about 10 000 8C) at localized spots on the electrodes such that the electrode materials will be vaporized, leaving craters behind on the surface as evidence of material removal [1]. Research has proven that the following material removing mechanisms are acting during EDM [3]: · Material removal through thermal shock. · Chipping of re-solidified material. · Material removal through melting and ejection processes. · Dissolution of grains by removing phases of higher electric conductivity. The uneven material removal at the anode (tool) and the cathode (workpiece) due to the uneven distribution of the discharged energy allows the economical use of spark erosion. Some institutes in Germany and Japan are conducting research on the fundamentals of the EDM process and optimization of the machines. The WZL (Aachen), for example, is committed to basic research on erosion mechanisms and process modeling, surface and rim zone formation, EDM of innovative materials (e.g. ceramics), workpiece behavior and fatigue strength [4, 5]. The IPT (Dresden) is working with special materials for EDM such as tungsten carbide and electroconductive ceramics and is developing the necessary technology for these materials in SEDM and WEDM. Another research topic is the analysis of single discharges and the influence of contamination in the gap on these single discharges. Together with the IGET (Magdeburg), the IPT optimized the influence of the current pulses on the process during the machining of tungsten carbide with WEDM. The main research topic of the IGET in the field of EDM is the techno-

Fig. 7-1 Principle of EDM [2]

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logical exploitation of electrical pulses and discharges. Here research is directed towards EDM and ECM processes, focusing on pulse power technologies [electrohydraulic effect (EHE), electromagnetic effect (EME)], particle separation, modeling and analysis of energy conversion processes and electric, electronic and control equipment for application and opening up new fields of application [6–13]. The Department of Mechanical Systems Engineering at Tokyo University of Agriculture and Technology is researching in the field of gaseous dielectrics. Research is also concerned with fundamentals of the EDM process such as spark location prediction and the use of transistor-type isopulse generator and servo feed control for micro-EDM [14–18]. 7.1.2 Micro-EDM

Micro-EDM is a specific type of EDM. The term is used if the size of the product or structural details are smaller than about 500 lm. The machined geometries can be holes, cavities or pins. Typical machining technologies for microEDM are wire EDM (WEDM), sinking EDM (SEDM), EDM drilling, EDM milling and wire electrodischarge grinding (WEDG). Concerning the requirements for micro-EDM, some changes are necessary compared with conventional EDM to obtain sufficient accuracy and surface quality [19]: · The size of the craters must be small. · The discharge gap must be small (only a few microns). · The precision of movement must be high. · The precision of the tools (electrodes) and the clamping system (repeat accuracy) must be high (*1 lm). Micro-EDM has several advantages over cutting technologies: · Complicated contours are simple to manufacture in a single operation. · Geometric accuracy, without burring. · High reproducibility. · Non-contact procedure, hence no mechanical stresses. · Applicability to hardened workpieces and cemented carbide. · Production of structures with high aspect ratios. · Manufacturing of structures that cannot be produced with cutting processes (e.g. very small drill holes with diameter < 20 lm). For sinking EDM there are only a few machines on the market which are specially developed for micro-EDM. These special micromachines are characterized by small design (desktop machines), short distance between machine and generator (to avoid the loss of energy in the pulse), special generators that permit a pulse length of a few nanoseconds, axis movement accuracy < 1 lm, a resulting white layer (heat-affected area) < 1 lm and an incorporated WEDG unit to dress small electrodes. These specialized machines come from Switzerland (Sarix) and Japan (Panasonic).

7.1 Introduction

7.1.3 Die-Sinking EDM

Engraving and drilling are the two main application areas of die-sinking EDM (in its different forms). With classical engraving, the tool electrode is mapped on the workpiece through a forward feed movement in the z-direction. Hence the information on the workpiece geometry is in the tool (Fig. 7-2). One of the most important technologies in the field of mold insert production using wear-resistant materials and also in the production of microcomponents is die-sinking EDM. Structured electrodes with a positive geometry of the required structure are used for the sinking process to machine a cavity. The main challenge of die-sinking micro-EDM is the machining of structured electrodes. The cheapest way is machining by a cutting process. Referring to the milling process (see Chapter 5), the electrode size is limited to the smallest commercial milling tools. LIGA technology (see Chapter 8) permits the machining of electrodes with very small 2.5-dimensional structures with comparatively high costs for these electrodes [21, 22]. A new method is the use of laser sintering for the production of electrodes [23]. Another approach is the use of WEDM (see Section 7.1.4) to machine electrodes. By using this technology, the geometric variety is limited and the machining of sculptured surfaces is not possible. With WEDG (see Section 7.1.5), the machining of very small rods, squares, triangles or half cylinders becomes accessible. Small drillings with diameters of about 10 lm are only possible using WEDG electrodes. Research in Germany in this field is being carried out by the WZL (Aachen) on the analysis of strategies for tool and die making, high-power EDM, investigation of different electrode materials and working media, investigation of process implied emissions, gap width control and adaptive process optimization with artificial intelligence [5]. 7.1.4 Wire EDM

WEDM was introduced industrially in 1969, even though it had already been suggested as a process variant in 1949. Since then, this process has been estab-

Fig. 7-2 Die-sinking EDM [20]

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7 Micro-EDM for Mold Inserts Fig. 7-3 Wire EDM [20]

lished in manufacturing for the production of cylindrical and/or conical breaches. As the name suggests, wire EDM uses an unwinding wire (e.g. made from copper, brass, steel or molybdenum with a diameter 20–300 lm) that must constantly be renewed as a tool instead of a three-dimensional electrode (Fig. 7-3). In Germany it is again the WZL (Aachen) which is researching in this field. With thin wires (diameter 50 lm), the machining of microstructures with a width of < 10 lm, length 200 lm and an accuracy within ± 2 lm in HS6-5-3 (PM powder-metallurgical high-speed steel) and cemented carbide K40 (87.2% WC, 12% Co) was achieved [5, 24, 25]. 7.1.5 WEDG

The principle of WEDG was developed in the early 1980s [26]. WEDG is a special form of EDG (electrodischarge grinding). In this process, a moving wire between one or two guides is used as a tool electrode instead of a disk. The wire diameter is between 50 and 200 lm. The principle of a WEDG-unit is shown in Fig. 7-4. In this setup, the rotating workpiece is fed downwards and the discharge area is limited to the front edge of the wire. By using two wire guides, it

Fig. 7-4 Principle of WEDG

7.1 Introduction

is possible to make cutting operations between these guides. The advantage of one guide is higher repeatability and accuracy. WEDG is very flexible concerning the geometry of the machined structures in comparison with conventional EDG by using a block electrode. The use of a block electrode has several negative impacts on the process; for example, the wear has a direct influence on the accuracy of the rod. Another problem is the unevenness of the gap distance along the axis of the machined rod, which is influenced by the gradient of the concentration of the debris in the gap. If the variation in the debris concentration is large, the evenness of the gap distance will not be suitable for a stable and high-quality process. The development of CNC-controlled machines allows the machining of different shapes such as angles at the bottom of the workpiece and conical or spiral structures with the WEDG process. Research in this area is being performed by the Institute for Machine Tools and Factory Management (Berlin). They developed an EDM machining center for microtechnology based on a commercial machine. With this machine, a combination of the applications of WEDM, SEDM and EDG was realized. A main research topic is EDG. By using a structured disk for EDG, it was possible to machine channels with a thickness of about 100 lm and an aspect ratio of 20 [27–33]. At the international level, Japanese research institutes have long-standing experience in micro-EDM. The development of the WEDG unit at the Institute of Industrial Science (IIS, University of Tokyo) permitted the machining of very small and precise rods. The IIS is also developing other special applications such as the EDM lathe and EDM supported punching [26, 34]. These special applications will be considered later in this chapter. 7.1.6 EDM Milling

The EDM milling process is a promising technology in the field of EDM. With small electrodes with diameters of e.g. 50 lm (standard electrodes) layer by layer will be machined along defined paths comparable to the milling process. The main advantage is the omission of the structured sinking electrode. By eliminating the machining of the sinking electrodes considerable expense can be saved especially in the production of smallest batches and EDM then becomes more economical. The flexibility of the geometry is higher and small changes in the geometry can be made during the process. Although it has a few advantages, EDM milling is currently only rarely used in industry, for several reasons: · Missing CAD/CAM system for an EDM milling process. · There are few machines that are able to machine using several axes at the same time. · Difficult electrode wear compensation during the EDM process. · Inadequate machining strategy.

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At present, simple structures such as circles, squares or triangles can be machined. However, the main problem with EDM milling is the lack of an adequate CAD/CAM system. Without such a system, the realization of complicated structures such as gears with reasonable effort is impossible. The use of a CAD/CAM system from the milling process is difficult because in EDM there are different parameters and a special behavior in case of a short-circuit between the tool and workpiece. Electrode wear compensation is also a function to be implemented in the system: depending on the machined distance, an electrode feed in the z-direction is necessary to compensate the wear. To machine a real 3D-sculptured surface, multi-axis (more than two axes) machining is necessary. In this case the CAM system must make the backward movement in case of a short-circuit in the direction of the initial point along the machined line, otherwise the machined structure may be damaged. The electrode wear in the process has a significant influence on the geometric accuracy of the machined structure. There are several parameters which determine electrode wear: workpiece–electrode–material combination, diameter of the electrode, machining parameters (voltage, current, frequency, etc.), machining depth per line, etc. To reduce the influence of electrode wear, one strategy is the reduction of the machining depth per line to only a few microns. For each material–parameter combination some preliminary experiments have to be carried out to obtain some information about the wear-induced reduction of the tool length per machined distance. This information must be entered in the CAM system to achieve the wear compensation during the process. These experiments are also important to obtain information about the lateral gap, which in turn is important for the accuracy of the machined structure. A new electrode wear compensation system was developed at the Department of Mechanical Engineering, Division PMA, Katholieke Universiteit Leuven, Belgium. This involves real-time wear sensing in layer-by-layer milling EDM. A problem with continuous wear compensation is the case of a slot in the workpiece (Fig. 7-5). Inside the slot, the EDM process is interrupted because of the large distance between the electrode and workpiece material. Since there is no wear at the electrode, high inaccuracy of the geometry will occur if the feed in the z-direction is continuing. Optical measurement is not feasible owing to the machining configuration and the presence of the dielectric. A solution is provided by indirect wear sensing based on discharge pulse evaluation. For real-time wear sensing in EDM milling, a tool wear sensing correlation was adopted: the worn volume of the electrode during a certain time is assumed to be proportional to the number of effective (normal) discharges during this time. For an optimized process a combination of two methods is necessary. In full material, continuous wear compensation with downward movement in the z-direction is necessary; in the case of absent material, the wear sensing stops this downward movement until material is encountered again. The results showed that this combined method permits the machining of accurate geometries even in the case of missing material. EDM milling was used to produce a mold insert with three-dimen-

7.1 Introduction

Fig. 7-5 Workpiece with slot

sional free-form structures with a square length of about 800 lm and a depth of 400 lm. The necessary electrodes for EDM milling were machined by a WEDG process in different shapes. The smallest electrode diameter used was 30 lm [35–37]. Concerning the influence of the machining strategy on the geometric inaccuracy conditional on wear, some investigations were carried out at the Masuzawa Laboratory at the IIS, University of Tokyo. To machine a rectangular geometry four different sequences were used (Fig. 7-6). The advantage of this procedure is that no pre-experiments are necessary to obtain information about the electrode wear because no wear compensation in the process is made. By machining the sequence shown in Fig. 7-6, the geometric inaccuracy in the z-direction is reduced to an average value. An important influence for the accuracy is the machining depth per line. A large machining depth causes high geometric inaccuracy (high average value). To reduce the corner wear of the

Fig. 7-6 Sequences for machining a rectangular structure. Source: IIS

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Fig. 7-7 Sequences for machining a circular structure

electrode, the use of a squared electrode is advantageous. Other research has been performed with circular geometries (Fig. 7-7). Two machining strategies were analyzed: machining with several circles with different radii and machining of the outside circle and the inside area with parallel lines only [38, 39]. 7.1.7 Synopsis of Micro-EDM Technologies

By comparing different EDM technologies such as WEDM, SEDM and WEDG, each technology has its specific limits concerning the machining result. Some of these limits are shown in Table 7-1.

7.2 State-of-the-Art in the Industrial Environment

Micro-EDM machines have been developed in Switzerland (AGIE, Charmille, Sarix) and in Japan (e.g. Mitsubishi, Panasonic, Makino). These are die-sinking machines and wire EDM machines. The development of new generators that permit machining with very short and selected pulses is a challenge for these companies. The ongoing development of micro-EDM machines today provides thin wires with diameters of 20 lm for WEDM [41].

7.2 State-of-the-Art in the Industrial Environment Table 7-1 Limits of micro-EDM [40]

Parameter

WEDM

SEDM

WEDG

Geometric accuracy (lm) Positioning accuracy (lm) Arithmetic average peak-to-valley height, Ra (lm)

±1 £1 0.08–0.1

±1 £1 0.2–0.3

±1 £1 0.06–0.1

Minimum line width (lm): Ligament Channel

20–40 50–60

20–40 20–40

10 120

Maximum aspect ratio: Ligament Channel Hole Smallest internal radius (lm) Smallest electrodes (lm)

20–30 60–80 – 20 20

15–25 10–25 10–100 10 10

20–30 60–80 – – 5

Fig. 7-8 Spinneret tool with smallest line width of 60 lm. Source: Sarix SA, Losone

The industrial knowhow is well developed and specialist mold and die toolmakers have ample experience. The company Sarix (Switzerland) and other specialised companies have proprietary knowhow that is not available to the public. In addition to the development of new generators and machines, Sarix also develops the application technology for their custumers. Some examples of industrially used die-sinking operations are shown in the Figs. 7-8 and 7-9. The applications shown are only examples of the industrial state-of-the-art. They just highlight what is possible with the EDM technology available today.

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7 Micro-EDM for Mold Inserts Fig. 7-9 Automotive nozzle for diesel, diameter 8 lm. Source: Sarix SA, Losone

7.3 Mold Insert Production by Die-Sinking EDM

When the Institute of Production Science (wbk) started with EDM, a first challenge was the machining of a mold insert for a microwheel. Examples of milled electrodes and the associated mold inserts are shown in Fig. 7-10. When using several identical electrodes, it is important to use a highly precise clamping system with a repeating accuracy of only a few microns [42]. After optimization of the parameters and the number of electrodes, the mold insert for a gear housing (part of the demonstrator in the Collaborative Research Center SFB 499) was machined by die-sinking EDM. As the aspect ratio of this part is high (smallest cavity 100 lm, depth 600 lm) it was not possible to machine

Fig. 7-10 Electrodes and moldinserts for a wheel

7.4 Mold Insert Production by EDM Milling Fig. 7-11 Mold insert of a gear housing, part of the SFB 499

this mold insert by a cheaper and faster milling operation. Instead of one complex electrode, five aligned electrodes were used: four very simply shaped turning parts and one electrode for the gear which was milled. In Fig. 7-11 the resulting mold insert is shown. In the next stage in the research project, this part will be machined by an EDM milling process.

7.4 Mold Insert Production by EDM Milling

Today, EDM milling is not just a topic in research institutes but also in manufacturing companies. One example is the work done by the company Sarix. They developed a micro-EDM machine for EDM milling which has six axes (including a controlled rotation axis) and permits machining along four axes at the same time. In a subsequent step they adapted a CAM system from conventional milling applications and modified it for the special requirements of EDM milling. By using a WEDG unit with one guide (see Section 7.1.5), the machining of small electrodes with diameters of 10 lm is possible. Some results for EDM-milled structures obtained by Sarix are shown in Fig. 7-12 [43]. The wbk is also researching in the field of EDM milling. In these investigations, two directions are being analyzed: EDM milling with standard electrodes made of metal and EDM milling with a new type of electrode developed at the wbk. In the field of standard electrodes, the research topic at the wbk is the optimization of EDM milling parameters concerning infeed, machining time and surface roughness for different electrode diameters and material combinations. The advantage of using standard electrodes is the small diameter of the tool electrode. Commercial electrodes made of cemented carbide (WC–Co) are available with diameters down to 50 lm. By using a WEDG process, electrodes can be machined with diameters down to 10 lm (aspect ratio 30). For some applications (e.g. finishing), it is necessary to use these electrodes for EDM milling. For rough machining, the use of the newly developed electrode might be useful. At the moment the necessary pre-experiments are being carried out at the wbk in order to machine mold inserts with EDM milling in the near future.

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7 Micro-EDM for Mold Inserts Fig. 7-12 EDM-milled structures. Source: Sarix SA, Losone

7.5 Development of a New Type of Electrode for EDM Milling

As already mentioned, the main problem in EDM milling is the electrode wear compensation. In order to avoid the pre-experiments concerning the electrode wear, a new type of electrode with a wear compensation that is independent of the actual electrode wear has been developed at the wbk. The principle of the new type of electrode is the following. A non-conductive capillary is used to carry a conductive medium (e.g. paste or liquid). The conductive material is the real electrode that implements the material removal. Independently from the conductive material consumption, permanent replenishment and the removal of the excess electroconductive material permit a constant length of the electrode. For the realization of small ‘electrode’ diameters, the use of an electroconductive liquid is advantageous. In order to eliminate the removed material and the excess fluid from the working area, the electrode was put on the machine table and the workpiece was placed above the electrode. This design is different from the conventional EDM configuration. The principle of this electrode is shown in Fig. 7-13. The function of this design is as follows. An aluminum collet is wrapped around the capillary, and this collet is fixed with a gasket on to the construction. All elements from the setup are made of stainless steel. They connect the nonconductive capillary through a valve with the reservoir containing the fluid. The whole system is put in an impounding basin made of stainless steel, which is connected with a squeeze pump that feeds the fluid back into the reservoir. The weight of the fluid allows a continuous flow and the flow rate is controlled by the valve with scale. The flow rate depends on the diameter of the capillary used. If the inner diameter of the capillary becomes < 1.5 mm, the application of additional pressure is necessary. This is realized by charging the compensating reservoir with compressed air. The narrow but high reservoir

7.5 Development of a New Type of Electrode for EDM Milling Fig. 7-13 Principle of the new electrode

used permits a high fluid column pressure. The electric pulse is applied by a bolted connection close to the collet. As already mentioned, the use of a fluid as electrode is advantageous in order to realize small diameters. The application of a non-conductive capillary made of ceramics permits the carriage of the fluid to the process area and allows for different diameters. The selected capillaries have the measurements given in Table 7-2. These capillaries are made of aluminum oxide (Al2O3) possesing a resistivity of 1014 X cm at a temperature of 20 8C. The choice of the fluids was more complex in comparison with the choice of the capillaries. In chemistry all acids and bases are conductive. Mercury is a liquid metal but it is poisonous. Electroconductive chemical compounds are saline solutions that are also used in ECM (electrochemical machining) as electrolytes. Many different salts are known in chemistry. To avoid any danger for the operator and machine, only salt compounds were used that have no risk and safety problems. In Table 7-3 the selected compounds are listed along with their electrical conductivities for saturated solutions. First experiments were conducted with sodium chloride fluid combined with a ceramic capillary (outer diameter 3 mm). The experiments were carried out on a micro-EDM machine at Sarix. The parameters used were the following: · frequency: 90–150 kHz; · voltage: 200–250 V; · current: 30 A; · gap: 0.1 mm; · workpiece material: aluminium, steel. Table 7-2 Selected capillaries made of aluminum oxide (Al2O3)

Outer diameter (mm) Inner diameter (mm) Wall thickness (mm)

0.5 0.2 0.15

0.6 0.3 0.15

1 0.5 0.25

2 1 0.5

2.5 1.3 0.6

3 2 0.5

4 3 0.5

5 4 0.5

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7 Micro-EDM for Mold Inserts Table 7-3 Selected electrolytes

Name

Formula

Electrical conductivity (S/m)

Sodium chloride Potassium chloride Potassium bromide Ammonium iodide Potassium iodide

NaCl KCl KBr NH4I KI

21.51 28.10 35.07 42.00 42.26

The process is a mixture of EDM and ECM. Concerning the electrical field, a ‘real’ EDM process with a working gap was not possible. Even if the distance between the liquid and workpiece was 0.1 mm, the electrical force at a voltage of 250 V was very high and the liquid was pulled to the workpiece. The physical contact combined with the electrical circuit from the EDM generator starts an ECM process. As a reaction product, a gas is generated at the contact area between the liquid and the workpiece. Gas sparks were observed in that area which are responsible for the main material removal as the machined surface looks like an EDM surface with many craters and not like an ECM surface. For diameters s < 0.5 mm, it is expected that the electrical conductivity of the medium inside the ceramic capillary might be too small to carry the energy of the discharge. In this case, the use of a capillary made of metal (e.g. cemented carbide) is necessary. As the distance between the capillary and the workpiece is very large (about 1–2 mm), there is no discharge in the gap between the metal capillary and the workpiece. Since in the EDM process the electrode wear is fairly small and the process is a combination between ECM and EDM, no wear problems are expected at the metal capillary. However, so far no experiments with these metal capillaries have been conducted with this special electrode setup. Such investigations are scheduled for the near future.

7.6 Special WEDG Applications for Micro-EDM

The main application for WEDG is the production of small pins that serve in most cases as electrodes for the machining of very small holes. The WEDG with a guide shown in Fig. 7-4 permits the machining of rods with diameters < 20 lm. In Fig. 7-14 an electrode is shown with a diameter of 12 lm and a length of about 400 lm [44]. With a special generator for micro-EDM, it is possible to machine rods with a diameter of a few micrometers. The machining setup for the example in Fig. 7-10 was the following: frequency = 170 kHz, voltage = 80 V and capacitance = stray capacitance (no added capacitors). The use of the SX-MPS generator of the Sarix EDM machine allows for working pulses with a length of a few nanoseconds. With these short pulses it is possible to reduce the heat-affected zone

7.6 Special WEDG Applications for Micro-EDM Fig. 7-14 Rod machined by WEDG

(white layer) to almost zero. The machined surface has roughness values Ra = 0.1 lm and Rz = 0.6 lm. An application of WEDG to mold inserts can be found in the structuring of cores and very small ejection pins. With a numerically controlled rotation axis the machining of a helically structured core is also possible. The machining of cores will be a research topic for the next project period. 7.6.1 EDM of Milling Tools

As already mentioned in Chapter 5, the wbk is also researching in the field of micromilling of hardened steel. To explore a new way to machine milling tools, the adequacy of the WEDG process for this task was analyzed. There are several advantages to using the EDM process for the machining of milling tools, such as easy variation of the geometry and the very small process forces that avoid the development of stresses in the workpiece [45]. Together with the IIS (University of Tokyo), the wbk has carried out first experiments to machine milling tools with EDM. The diameter of these tools was set to 100 lm to allow for a comparison with conventional ground tools. The milling process demands a clearance area at the back side of the milling tool to avoid additional stress on the tool. As round geometries can tolerate about 30% higher forces in comparison with trapezoidal geometries, a semi-circular cross-section would be the best choice. Owing to restrictions of EDM machining, the semi-circular shape was approximated by a polygon with 36 corners. To reduce the diameter at the back side, a WEDG process in 18 steps with angles of 108 was used. The tool has two clearance angles (main and side clearance angles), both 108. The tool shown in Fig. 7-15 was machined on the EDM machine at the IIS (Panasonic MG ED72W). The wire on the WEDG unit was made of brass with a diameter of 100 lm and the feed rate was about 15 mm/min. The parameters used for the finishing were voltage = 80 V, capacitance = stray capacitance (no

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Fig. 7-15 EDM-machined milling tools made of cemented carbide, diameter 100 lm. Source: wbk and IIS

added capacitors) and machining feed = 2 lm/s. The resulting surface roughness values were Ra = 0.06 lm and Rz = 0.1 lm. The tool was machined in five main steps: cutting the bottom, reducing the diameter from the initial dimension of 0.3 mm to 0.1 mm, machining the half cylinder in two steps, machining the clearance area at the back side of the tool in 18 steps and machining the clearance angle at the end of the milling tool. The machining time for the tool was about 45 min. The usability of the EDM-machined milling tools was demonstrated by some milling experiments. These were performed on the EDM machine at the IIS even though the machine does not have ideal characteristics for the milling process. The advantage of using this machine was the prevention of inaccuracy by repeated chucking. In Fig. 7-16, some milling results are shown for a workpiece made of brass. As the lines show, it was a real cutting process with the tool. The tool’s cutting edge had sufficient sharpness for the milling process. The parameters for the cutting experiments were the following: · feed per cutting edge = 3 lm; · feed rate = 6 mm/min;

Fig. 7-16 Groove produced by milling using a WEDG tool. Source: wbk and IIS

7.6 Special WEDG Applications for Micro-EDM

Fig. 7-17 Milling burr. Source: wbk and IIS

· rotation = 2000 rpm; · depth of cut = 10 lm; · machining length = 3–15 mm. The experiments showed an astonishing behavior of the tool. Instead of fracture which was expected in view of the high process force, an excursion of the spindle was observed. The rigidity of the spindle was too low for the milling operation. Another problem is the strong burr at the groove, as can be seen in Fig. 7-17, upper pictures. There are several possible reasons for this burr: the deficient tool geometry regarding the clearance angles and as the main problem the low cutting speed. The latter arose because the experiments were carried out on an EDM machine with a belt collet. 7.6.2 EDM Lathe for Micromachining

The machining of small rotational symmetrical turning parts (diameter < 1 mm) in hard materials is a challenge for machining technologies. The machining of rods can be done by WEDG but in this case it is not possible to machine sharp corners. Some parts may have microdrillings inside and microstructures such as channels or different small diameters outside. For this application, the IIS identified the necessity for the development of an EDM lathe machine. In Fig. 7-18 both main machining systems for EDM turning are shown. Regarding the wire electrode, the principle is the same as in WEDG. Between the running wire and the rotating workpiece the EDM process reduces the diameter of the workpiece. If a crank is machined with the wire electrode, the passage between different diameters has as a minimum radius the electode radius. If a sharp corner is required, the use of a disK electrode is preferred (conventional EDG). In this case, the disk electrode and the workpiece are rotating and the EDM process takes place between the two rotating parts. For this machining setup two parallel spindles are necessary. The larger the diameter of the

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7 Micro-EDM for Mold Inserts Fig. 7-18 Principle of EDM turning. Left: EDM turning with a wire electrode. Right: EDM turning with a disk electrode. Source: IIS

disk electrode, the sharper is the machined corner because the electrode edge used is very long and the wear at the corner is low. The workpiece can be the component to be machined but it can also be the electrode for the drilling operation. If a small disk electrode is machined by EDG, the machining of inner turning geometries such as the seat of a rolling bearing is possible. Since all machining steps are performed in one chucking, the production tolerance is very high [46]. In Figs. 7-19 and 7-20 some examples of EDM-turned parts are shown. 7.6.3 Micropunching with WEDG Support

By micropunching, microholes can be machined efficiently. However, the problems in realizing practical applications of this technology are the following: · difficulties in machining the small punch; · difficulties in machining the small hole (die); · difficulties in assembling the punch and the die. The IIS identified the possibility of solving these problems by applying the WEDG unit on a punching machine. By using the WEDG at several stages of the punch and die machining process, these problems can be solved.

Fig. 7-19 Workpiece (left) and electrode after EDM turning of a groove (right). Workpiece material, stainless steel; electrode material, WC alloy. Source: IIS

7.6 Special WEDG Applications for Micro-EDM Fig. 7-20 Stepped cavity, sectional view. Source: IIS

The procedure leading to this punching system is divided into four parts (Fig. 7-21): 1. The electrode for machining the die is made by WEDG. 2. The die is machined with an EDM process by using this electrode. 3. The punch is machined by WEDG, removing the thinner part of the electrode used. The tool assembly is complete. 4. The workpiece is punched. The three difficulties mentioned above are eliminated by this system because: · WEDG can machine a very thin rod down to a diameter of 10 lm. This makes the production of a small punch easy. · WEDG can also machine a very thin electrode for EDM. The machining of a small hole or the die is not difficult since the associated machining technology is developed.

Fig. 7-21 Principle of WEDG-supported punching. Source: IIS

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7 Micro-EDM for Mold Inserts Fig. 7-22 Punched holes. Source: IIS

· A special advantage of this system is that the center lines of the electrode, the die and the punch are common throughout the whole process. There is no need to adjust the centers, which was a major challenge in completing the punch and die assembly. In the Fig. 7-22 an example of punched holes is shown. By using a controlled rotation axis for the WEDG process, the machining of angular geometries such as irregularly shaped holes is possible [47].

7.7 Conclusion

Among the different machining methods used for micromachining, EDM technology is an excellent choice for the production of tiny structures in a variety of materials. The principle of machining is the repeated removal of very small parts of the workpiece. Micro-EDM is characterized by high accuracy, low speed and shape flexibility for use with all metals and alloys. This characteristic is very promising for the machining of tools, dies and molds and the production of prototypes or small series of microparts. This chapter has presented vatrious applications of micro-EDM. A popular application is die-sinking EDM. This technology is used for the machining of dies and molds for microstructures. The economic production of the electrodes, usually done by milling, sets the limitation on further miniaturization. To avoid these microstructured electrodes, EDM milling is a potential alternative but so far the boundary conditions for a successful application such as a CAM system or the machining strategies are still lacking. The development of a new type of electrode for EDM milling describedabove might help to avoid the geometric inaccuracy resulting from electrode wear in the EDM milling process. However, at present this is still a research topic and there some further development is necessary to achieve a serviceable tool for EDM milling. In view of the growing competition between micromilling, laser beam machining and micro-EDM, new application fields for micro-EDM are being ex-

7.8 References

plored. Some research on the enhancement of applications for WEDG is in progress in Japan. Technologies such EDM lathe and WEDG punching have been developed in recentyears. A special application is the machining of micromilling tools by WEDG which is under investigation at the wbk. First results of machining these tools are promising. The tendency towards miniaturization will continue in the future, increasing the importance of micro-EDM technology in various fields of production. The exploration of new technologies and applications and the process-sure use of EDM milling are important elements of this development and, hence, will be research topics at the wbk and the IIS in the future.

7.8 References 1 D. Allen, in: Microelectrodischarge Ma-

2 3

4

5 6

7 8

9

chining, Micromachining of Engineering Materials, J. McGeough (ed.); New York: Marcel Dekker, 2002, pp. 179–201. W. König, Abtragen, 3rd edn., Berlin: Springer Verlag, 1997. M. Timm, Elektronische Stromquellen für das funkenerosive Schneiden von elektrisch schlecht leitfähigen Werkstoffen; Dissertation, Otto-von-Guericke-Universität Magdeburg, 1996. F. Klocke, D. Lung, G. Antonoglou, D. Thomaidis, The effects of powder suspended dielectrics on the thermal influenced zone by electrodischarge machining with small discharge energies. In: Proceedings of ISEM XIV; 2004. WZL homepage, www.wzl.rwthaachen.de (accessed August 2004). K. Künanz, H. Juhr, E. Knösel, R. Haas, Prozessnahe Einzelentladung zur Analyse des Funkenerodierens. Werkstattstechnik 11, 1998, 11–13. IGET homepage, www.uni-magdeburg.deiget (accessed August 2004). W. Rehbein, H.-P. Schulze, K. Mecke, G. Wollenberg, M. Storr, Influence of selected groups of additives on breakdown in EDM sinking. In: Proceedings of ISEM XIV; 2004. H.-P. Schulze, M. Läuter, G. Wollenberg, M. Storr, W. Rehbein, Investigation of the pre-ignition stage in EDM. In: Proceedings of ISEM XIII; 2001, pp. 141–152.

10 H.-P. Schulze, G. Wollenberg, M. Läuter,

11

12

13

14

15

16

17

H.-J. Trautmann, M. Zybura-Skrabalak, Structure of process energy sources for timeparallel combined processes. In: Proceedings of ISEM XIV; 2004. H.-P. Schulze, R. Herms, H. Juhr, W. Schatzing, G. Wollenberg, Comparison of measured and simulated crater morphology for EDM. In: Proceedings of ISEM XIV; 2004. G. Wollenberg, H.-P. Schulze, Th. Pape, Process parameter and analysis for an optimal feed rate control of EDM. In: 16th International Conference on Computer-aided Production Engineering, CAPE 2000, Edinburgh, ImechE Conference Transactions; 2000, pp. 183–190. G. Wollenberg, H.-P. Schulze, Th. Pape, M. Läuter, Moderne Generatortechnik für die funkenerosive Bearbeitung. In: 4. Aachener Fachtagung Funkenerosive Bearbeitung, November 1999, Aachen; 1999, pp. 91–26. M. Kunieda, S. Fujuoya, Improvement of EDM efficiency by supplying oxygen gas into gap. Ann. CIRP 1991, 40, 215–218. M. Kunieda, M. Yoshida, Electrical discharge machining in gas. Ann. CIRP 1997, 46, 143–146. M. Kunieda, C. Furudate, High precision finish cutting by dry WEDM. Ann. CIRP 2001, 50, 121–124. M. Kunieda, Y. Miyoshi, T. Takaya, N. Nakajima, Z.-B. Yu, M. Yoshida, High

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18

19 20

21

22

23

24

25

26

27

28

29

30

speed 3D milling by dry EDM. Ann. CIRP 2003, 52, 147–150. M. Kunieda, F. Han, S. Wachi, Improvement of machining characteristics of micro-EDM using transistor type isopulse generator and servo feed control. Precision Eng. 2004, 28, 378–385. T. Masuzawa, Micro-EDM. In: Proceedings of ISEM XIII; 2001, Vol. 1, pp. 3–19. W. Schweizer, in: Moderne Fertigungsverfahren der Feinwerktechnik: ein Überblick, W. Schweizer, L. Kiesewetter (eds.); Berlin: Springer, 1981, 25–34. A. Wolf, W. Ehrfeld, H. P. Gruber, Mikrofunkenerosion für den Präzisionsformenbau. wt Werkstattstechnik 1999, 89, 499 –502. A. Wolf, W. Ehrfeld, H. Lehr, F. Michel, M. Nienhaus, Combining LIGA and electrodischarge machining for the generation of complex microstructures in hard materials. In: Proceedings of 9-IPESUME4, Braunschweig; 1997, pp. 657– 660. J. P. Kruth, T. Laoui, Selective laser sintering: state of the art. Int. J. Electr. Mach. IJEM 2001, (6), 7–17. F. Klocke, D. Lung, T. Nöthe, Microcontouring by EDM with fine wires. In: Proceedings of ISEM XIII; 2001, pp. 767–778. F. Klocke, D. Lung, D. Thomaidis, G. Antonoglou, Using ultra thin electrodes to produce microparts with Wire EDM. In: Proceedings of ISEM XIV; 2004. T. Masuzawa, M. Fujino, K. Kobayasi, T. Suzuki, Wire electro-discharge grinding for micromachining. Ann. CIRP 1985, 34, 431–434. G. Spur, E. Uhlmann, U. Doll, N. Daus, WEDM of micorstructured component parts – heat conduction model, Int. J. Electr. Mach. IJEM 1999, (4), 41–46. S. Thiel, S. Lehnicke, O. Zimmer, D. Grimme, U. Doll, S. Piltz, Abformung von Mikrostrukturen in Glas. VDI-Z VII; November 2000, pp. 26–30. E. Uhlmann, U. Doll, Mikrofunkenerosion zur Fertigung von Mikrosystemen. In: Jahrbuch Innovative Produktion 2000; Berlin: Berlin Trade und Contact Verlag, 2000. E. Uhlmann, U. Doll, S.Piltz, Electrical discharge grinding of microstructures.

31

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33

34

35

36

37

38

39

40

41

Int. J. Electr. Mach. IJEM 2001, (6), 41– 46. E. Uhlmann, U. Doll, S. Piltz, Electrical discharge grinding (EDG) using microstructured disk electrodes. Ann. WGP 2001, 8, 74–81. E. Uhlmann, S. Piltz, S. Jerzembeck, Micromachining of cylindical parts by electrical discharge grinding. In: Proceedings of ISEM XIV; 2004. E. Uhlmann, U. Doll, Application of l EDM in the machining of microstructured forming tools. Presented at the 3rd International Machining and Grinding Conference, Cincinnati, OH; 1999. T. Masuzawa, State of the art of micromachining. Ann. CIRP 2000, 49, 473– 488. D. Reynaerts, W. Meeusen, H. Van Brussel, Machining of three dimensional microstructures in silicon by electrodischarge machining. Sens. Actuators A 1998, 67, 159–165. W. Meeusen, D. Reynaerts, J. Peirs, H. Van Brussel, V. Diericks, W. Driesen, The machining of freeform micromolds by micro-EDM; work in progress. In: Proceedings of MME 2001 (Micromechanics Europe Workshop), Cork, Ireland; 2001, pp. 46–49. P. Bleys, J.-P. Kruth, B. Lauwers, Sensing and compensation of tool wear in milling EDM. In: Proceedings of ISEM XIV; 2004. T. Masuzawa, Y. Zuyuan, M. Fujino, 3D micro-EDM with simple shape electrode. Int. J. Electr. Mach. IJEM 1998, (3), 7–12. T. Masuzawa, Y. Zuyuan, M. Fujino, 3D micro-EDM with simple shape electrode, Int. J. Electr. Mach. IJEM 1998, (3), 71– 78. J. Fleischer, J. Schmidt, M. Knoll, S. Haupt, C. Müller, R. Förster, A. Gehringer, Mikrobearbeitung durch Abtragen. wt Werkstattstechnik 2004, 94, 589. J. Hesselbach; A. Raatz, J. Wrege, H. Hermann, A. Wenda, H. Weule, C. Buchholz, H. Tritschler, M. Simon, F. Klocke, M. Weck, A. v. Klitzing, J. v. Bodenhausen, In: MikroPRO – Untersuchung zum internationalen Stand der Mikroproduktionstechnik, J. Hesselbach (ed.), Schriftenreihe des Instituts für Werkzeugmaschinen und Fertigungs-

7.8 References technik der TU Braunschweig; Essen: Vulkan Verlag, 2002, pp. 134–139. 42 J. Schmidt, J. Fleischer, M. Knoll, Electrodes for Micro-EDM. In: Proceedings of EUSPEN International Topical Conference on Precision Engineering, MicroTechnology, Measurement Techniques and Equipment; 2003, Vol. 1, pp. 177–179. 43 J. Fleischer, J. Schmidt, M. Knoll, E. Salim, F. Leleu, Zündender Funke für die Mikrotechnik – Anwendungen und Maschinenkonzepte der Mikrofunkenerosion. Mikroproduktion 2004, (1), 36–38. 44 J. Fleischer, T. Masuzawa, J. Schmidt, M. Knoll, New applications for micro-

EDM. In: Proceedings of EUSPEN; 2004, pp. 43–44. 45 J. Fleischer, T. Masuzawa, J.Schmidt, M. Knoll, New applications for micro-EDM. J. Mater. Process. Technol. 2004, 149, 246– 249. 46 T. Masuzawa, K. Okajima, M. Fujino, EDM lathe for micromachining. Ann. CIRP 2002, 51, 355–358. 47 T. Masuzawa, M. Yamamoto, M. Fujino, A micropunching system using wireEDG. In: Proceedings of ISEM 9; 1989, pp. 86–89.

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8 Lithographic Fabrication of Mold Inserts M. Guttmann, J. Schulz, V. Saile, Institute of Microstructure Technology, Forschungszentrum Karlsruhe, Germany Abstract The fabrication process for X-ray LIGA mold inserts is reviewed. The individual steps in the fabrication sequence are described in detail including their technical background. Specifications for mold inserts are presented. Physical limitations such as secondary radiation or the swelling of the PMMA resist in aqueous electroplating solutions are discussed. Finally, opportunities and difficulties are summarized for an alternative fabrication process, namely UV lithography with SU-8 resist.

Keywords LIGA; lithography; electroforming; mold insert fabrication; specifications; reliability; quality standards

8.1 8.2 8.2.1 8.2.2 8.2.3

Introduction 188 The LIGA Process for Mold Inserts Process Overview 190

190

8.2.4 8.2.5 8.2.6 8.2.7 8.2.8 8.2.9 8.2.10

Technology for Mask Fabrication 192 Substrate Issues and Resist Coating for Mold Inserts 193 Exposure Process 196 Development Process 199 Electroforming of the Mold Insert 201 Machining the Mold Insert to its Final Contour 205 Polishing the Mold Insert 206 Cleaning of the Mold Insert 207 Mold Insert Specifications and Their Control 210

8.3 8.3.1 8.3.2 8.3.3

Precision-limiting Effects 211 Future Goals to Advance LIGA Mold Insert Fabrication 213 Alternative to X-Ray LIGA: UV Lithography 214

Advanced Issues in LIGA Fabrication

211

Advanced Micro and Nanosystems Vol. 3. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31208-0

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8.4 8.5 8.6

Conclusions 215 Acknowledgment 217 References 217

8.1 Introduction

The LIGA process [1] was invented in the 1980s and since then has been further developed at the Institute of Microstructure Technology (IMT) at the Forschungszentrum Karlsruhe to prove its flexibility with respect to various applications and, in recent years, towards a reliable manufacturing technology. LIGA allows the production of microcomponents with complex lateral geometries, with resolution in the micron and sub-micron range and with structure heights up to the millimeter range. Typical materials employed are polymers, metals and ceramics, thus covering a wide range of ‘non-silicon’ candidates. The main process steps have given the technique its name LIGA, a German acronym for Lithografie (lithography), Galvanoformung (electroforming) and Abformung (replication techniques such as molding or hot embossing). It is worth noting that until recently in the USA replication was not considered an integral part of LIGA (Lithografie and GAlvanoformung). The most characteristic and unique step of LIGA is X-ray lithography using highly collimated and energetic X-rays from a synchrotron radiation source. Such X-rays penetrate hundreds of microns of polymer resist and hence can pattern it with extremely sharp, smooth and vertical sidewalls. In addition, owing to the short wavelengths of the X-rays, the spatial resolution of the process is fairly high, thus allowing for extremely high aspect ratios, i.e. the ratio of height to width of a structure. The first steps in the LIGA process sequence include fabrication of X-ray masks, typically with an electron-beam writer and subsequent Au electroplating of absorber structures, deposition of thick resist films up to several millimeters on a substrate and exposure to X-rays followed by development. Then, depending on the material and number of parts selected for the final product, different fabrication routes can be chosen, which may include further steps of microreplication through electroforming and/or a variety of molding techniques (injection molding, embossing, casting, compression molding, etc.). Therefore, the X-ray printed polymeric microstructure can be used: · Simply as it is (so-called ‘direct LIGA’) [2, 3]. · As a preform for ceramic microstructures if ceramic filled polymers or preceramic polymers are used [4]. · As an electroplating template to generate metallic microparts, which are often further processed by precision lapping to control thickness [5]. · As an electroplating template to produce a metallic master mold (mold insert), which can then be used multiple times to mold cost-effective replicates in other materials, primarily polymers [6].

8.1 Introduction Table 8-1 Aspect ratios achieved in deep X-ray lithography

Aspect ratio

Min. lateral dimensions (lm)

650

0.2

250 40 30

Structure height (lm)

Type (see text)

Microstructure

Reference product

130

A

Grating structure

2 Diameter = 26

500 1100

A B

100/100

3000

B

Grating structure Columns, here in SU-8 Lines and spaces

UV–VIS spectrometer (Fig. 8-1 a and b) IR spectrometer Columns (Fig. 8-1 c) Lead frames (Fig. 8-1 d)

As a consequence of the high energy and the parallel nature of X-rays from a synchrotron, very high structures can be produced with nearly vertical and extremely smooth sidewalls [7]. Examples of aspect ratios achieved are given in Table 8-1, where a distinction is made between a component (A) which is relatively large (mm to cm) but contains very fine (lm to nm) details (e.g. the diffraction grating of the LIGA microspectrometer [8–10]) and a free-standing microcomponent (B) with lateral dimensions in the lm range (e.g. for actuators, inertial sensors or optical components). Fig. 8-1 show relevant microstructures as examples for each case. The LIGA technology allows for [11]: · Large to very large structural heights (typically from hundreds to thousands of micrometers thick that can be formed in a single step). · Smallest lateral dimension (a few micrometers with structural details in the sub-micrometer range). · Roughness of the sidewalls below 30 nm, thus qualifying the sidewalls to be used as functional elements in microoptical systems (MOEMS). · Access to a large base of functional materials (electroplated metals and alloys, molded polymers, ceramics, composites, multilayered materials, graded materials and nanomaterials) [12]. · Formation of complex shapes. · Structural accuracy of features. · Mixing of scales (small features on large parts). · High-volume, low-cost production (the original concept of the LIGA process included molding, by injection molding or hot embossing, as the technology for mass fabrication). For a full exploitation of the LIGA potential, fabrication lines for LIGA components have been established complying with rigorous quality management systems [13].

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a

b

c

d

Fig. 8-1 (a) Microprecision component type A (see also Table 8-1): grating of a microspectrometer; (b) close-up SEM view of the diffraction grating; (c) microprecision component type B (see also Table 8-1): FZK-Logo with 5 lm thin walls; (d) microprecision component type B (see also Table 8-1): columns in SU-8

8.2 The LIGA Process for Mold Inserts 8.2.1 Process Overview

In deep X-ray lithography, the pattern on an X-ray mask is lithographically transferred with synchrotron radiation into an X-ray-sensitive resist layer, typically poly(methyl methacrylate) (PMMA); see Fig. 8-2 a. Exposure to X-rays results in chain scission of the long polymer molecules and hence a reduced average molecular weight in the exposed areas. A developer will dissolve the smaller molecules in the irradiated areas while leaving the non-irradiated areas unchanged.

8.2 The LIGA Process for Mold Inserts a

b

c

d

Fig. 8-2 Schematic depiction of the LIGA process. (a) X-ray exposure of PMMA resist (with working mask); (b) electrodeposition of metal in a resist template; (c) molding (here: hot embossing) of a polymer using a metallic mold insert; (d) post-processing of molded component, here dicing

The resist template placed on an electrically conducting substrate is then filled electrochemically with a metal suitable for molding, typically Ni, NiFe or NiCo (see Fig. 8-2 b). The deposition of the metal is continued significantly beyond the resist height until a thick metal layer has developed above the resist, forming a continuous and solid metal base-plate for the fragile microstructures (Fig. 8-2c ). This approach leads to molds with excellent height control and, through the rigid base-plate, to molds that can withstand the mechanical loads during the subsequent molding process. After de-molding, the polymer components can be post-processed, e.g. by dicing (Fig. 8-2 d), or they can be filled with metal by electrodeposition ([14]; see also Chapter 14). LIGA involves a long process chain which, in many aspects, is comparable in complexity to semiconductor manufacturing. Hence, establishing LIGA as a reliable manufacturing technology requires stable individual process steps with high yield. Major efforts were therefore directed towards establishing well-controlled ‘standard processes’. Furthermore, suitable equipment such as X-ray

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scanners, megasonic development tanks and electroplating facilities were developed along with the equipment for hot embossing actually to make use of the mold inserts. In the following, IMT’s standard X-ray lithography process and its characteristics, based on PMMA as resist, are described in more detail. 8.2.2 Technology for Mask Fabrication

LIGA masks are M = 1 : 1 shadow masks, i.e., they show the same lateral dimensions as the imaged product. Classical chromium-quartz masks, the standard for contact or proximity UV lithography, cannot be used for X-ray lithography for two reasons: (i) the quartz substrate absorbs a significant fraction of the incoming irradiation and (ii) the thin chromium layer (typically 0.1 lm) does not generate sufficient contrast between the areas to be exposed and those not to be exposed. The X-ray absorption in materials is proportional to the atomic number Z to the third power, i.e. in order to assure a good contrast, an absorbing high-Z material with sufficient thickness needs to be patterned on to a low Xray absorption carrier, a thin plate or membrane of low-Z material. The combinations which are used at the Forschungszentrum Karlsruhe are either: · for resist heights between 200 and 600 lm: a titanium membrane (Z = 22) with a thickness of *3 lm, patterned with gold absorbers (Z = 79) with a thickness up to 25 lm [15], or · for resist heights in excess of 600 lm: a 500 lm thick beryllium wafer (Z = 4) with nominally 35 lm of gold absorber [16]. Gold absorbers of such a thickness can nowadays be patterned through optical lithography, for example in AZ9260 or SU-8 resist. However, the desired accuracies of the LIGA technique cannot be exploited with such masks since diffraction effects lead to sidewall smearing of the mask pattern and other detrimental effects. The high-precision masks at Forschungszentrum Karlsruhe are therefore fabricated in two steps, first the so-called ‘intermediate mask’ by electron-beam lithography and subsequently the ‘working mask’ by X-ray lithography. For the intermediate mask a 3.2 lm thick PMMA layer is spun on to a silicon substrate coated with 3 lm of titanium. The central area of the silicon carrier is coated with a 100 nm thin carbon layer to prevent adhesion of the titanium to the silicon. The PMMA is patterned by a 100 keV electron-beam writer, thus generating the master pattern as indicated in Fig. 8-3 a. The use of 100 keV electrons minimizes scattering within the resist so that the smallest feature details can be significantly less than 1 lm with smooth sidewalls [17]. The developed PMMA is filled with 2.3 lm of gold by electroplating and the PMMA is subsequently stripped. Then, a carrier frame is bonded to the substrate surrounding the area with the microstructures. Because of the deliberately chosen low adhesion of carbon to the silicon substrate, the frame with the membrane can easily be delaminated from the silicon substrate (Fig. 8-3 b).

8.2 The LIGA Process for Mold Inserts a

b

Fig. 8-3 Schematic depiction of the fabrication of an X-ray mask. (a) Writing of the master pattern (intermediate mask) with an electron beam writer; (b) separating the titanium membrane (working mask) from the substrate. The intermediate mask has 2.3 lm of electroplated gold as absorbers and the working mask 20–30 lm

The contrast of the intermediate mask is just sufficient to structure a PMMA layer of *60 lm or less thickness. Therefore, in a second step, the intermediate mask is copied into a 60 lm thick PMMA resist on another titanium membrane or on a beryllium wafer using X-ray lithography at a wavelength of *1 nm. The PMMA template is then again electroplated with gold. After stripping, the final result is the so-called working mask. It has 20–30 lm high electroplated gold absorbers sufficient for patterning resist thickness of several hundred lm to several mm at X-ray wavelengths around 0.2 nm, the standard exposure conditions for LIGA. 8.2.3 Substrate Issues and Resist Coating for Mold Inserts

The first step in resist technology is to apply a resist on to a substrate. Several types of substrates, e.g. 4-inch wafers (Si, SiO2 and Al2O3) or special copper plates, can be used with IMT’s scanner stations at the ANKA synchrotron at Forschungszentrum Karlsruhe. For fabrication of LIGA mold inserts, the following criteria must be met: · The substrates must have sufficient stiffness that neither internal stress of the thick electroplated layer nor thermal stress after cooling to room temperature will cause significant bending. This is achieved by employing copper substrates of 8 mm thickness (with outer dimensions of 84 ´ 54 mm). · The substrates must have a high planarity of not more than 15% of the overall allowed bending of the finished mold insert. This criterion is met with a planarity tolerance of ± 5 lm. · The substrate surface must provide sufficient adhesion strength to the resist. This is accomplished with a 3 lm thick titanium layer on top of the substrates. Oxidizing the surface of the titanium layer to TiOx in hot NaOH– H2O2 solution results in an increased surface roughness and good adhesion.

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· The substrate surface must be electrically conductive for carrying the current for electroplating. The combination of a copper substrate with a Ti–TiOx layer is an adequate solution. · The developer solution must not etch the substrate. This is achieved by complete galvanic coating of the substrate with nickel or gold prior to the deposition of the titanium layer. · Finally, depending on the choice of process for separating the substrate from the mold insert, certain mechanical criteria (if pulled off) or electrical criteria (if cut off by electrodischarge machining, EDM) must be met. Copper substrates are compatible with both criteria. The German LIGA centers (BESSY Berlin [18], IMM Mainz [19] and IMT Karlsruhe) use the standard resist PMMA as the X-ray-sensitive polymer layer. The negative tone EPON SU-8 resist is, in principle, also an option, in particular owing to its high sensitivity and associated reduced exposure time. However, stripping of SU-8 is problematic and currently still limits its use to low aspect ratio patterns for mold inserts. The resist technology at IMT is as follows: the resist layer is attached to a substrate through gluing of a prefabricated, commercially available PMMA sheet

a

b

c

Fig. 8-4 (a) Simple robot for gluing the resist on to the substrate; (b) detail of robot with reservoir for PMMA sheets and holder for 4-inch wafers; (c) simple press to obtain a minimal gluing zone of just 10 lm

8.2 The LIGA Process for Mold Inserts a

b

Fig. 8-5 (a) PMMA resist sheet (63 ´ 23 ´ 0.4 mm; GS 233; unpolished); (b) avoiding manual gluing of resist sheets results in improved position control and minimal glue overflow

(GS233 from Röhm GmbH [20], referred to as ‘non-cross-linked PMMA’) to the substrate. Prior to gluing, the prefabricated PMMA sheets (see Fig. 8-5 a) are cut into rectangular or round pieces and then milled to the desired height (several hundred micrometers to several millimeters) with high precision (tolerance < 10 lm) and finally annealed under controlled conditions (heating at 15 8C/h to 110 8C, 1 h dwell at 110 8C and cooling at 8 8C/h to room temperature). The glue consists of PMMA dissolved in MMA (methyl methacrylate), mixed with DBPO (dibenzoyl peroxide) and DMA (dimethylaniline) and is applied by a robotic dispenser on to the substrate. The resist sheet is then transferred on to the substrate by an automatic pick and place machine (Fig. 8-4 a and b) and finally pressed on to it for several hours (Fig. 8-4 c). Advantages of this process route are a defined resist material quality, small variation in height, a glue zone of typically less than 15 lm thickness that behaves very similar to the resist under irradiation and subsequent development, good control of the position on the substrate (Fig. 8-5 b), good adhesion and finally a long shelf-life of the glued substrates. After gluing, in some cases (e.g. fabrication of mold inserts for optical applications) polishing of the resist surface is necessary to smooth out the grooves left from the milling process. A substrate-specific chuck is the only special part to perform this step on a standard polishing machine. The surface quality is significantly improved, as is demonstrated in Fig. 8-6. Alternative deposition technologies for the resist layer involve direct polymerization of a resin on the substrate or bonding a prefabricated PMMA sheet to an *3 lm thick resist layer spin-coated on to the substrate [21].

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Fig. 8-6 Improvement of surface quality through mechanical polishing of resist surface (see the microscopic and AFM pictures after the manual and the mechanical process)

8.2.4 Exposure Process

The LIGA technology requires a synchrotron beamline for resist exposure. Since the beam from the synchrotron forms a horizontal line only a few mm in height and much larger width in the exposure plane, a scanner is required for moving the mask–substrate package up and down (see Fig. 8-7 b) during exposure. The exposure dose calculation has to take into account that a broad spectrum of photon energies act on the mask–resist–substrate sandwich. The widely used program DoseSim developed at IMT [22, 23] calculates the exposure time for a given dose to be deposited at a specific synchrotron source and a beamline with optical components such as mirrors, windows and filters. At the ANKA synchrotron [24] with three beamlines (LIGA I, II and III; see Fig. 8-7 a) dedicated to three different resist height regions (as indicated in Table 8-2), standard exposure conditions have been established. To ensure complete development of the exposed domain of the resist, a minimum dose of 3.5 kJ/

Fig. 8-7 (a) Schematic depiction of one of the LIGA beamlines at ANKA Karlsruhe;

8.2 The LIGA Process for Mold Inserts 197

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8 Lithographic Fabrication of Mold Inserts

Fig. 8-7 (b) ANKA scanner at beamline LIGA II

Table 8-2 LIGA beamlines at ANKA and their targeted use

Beamline

Structure height (lm) Minimum width (lm) at aspect ratio 20 Structure details Verticality of PMMA structures (mrad)

X-ray lithography

Deep X-ray lithography

Ultradeep X-ray lithography

LIGA 1 Cr-mirror (cut off energy 4 keV) < 100 5

LIGA 2 Ni-mirror (cut off energy 6.9 keV) 100–600 5–20

LIGA 3 Full synchrotron spectrum (up to 40 keV) 600–3000 20–200

Sub-lm <1

Sub-lm 1.5

Not yet examined 1.5–2.5

cm3 for PMMA has to be deposited at the bottom of the resist. The dose at the top of the resist layer should not exceed 12–20 kJ/cm3 with higher top-dose values allowed for thinner resists. The top dose is limited by mechanical damage due to gas formation, the so-called foaming, of the resist surface. With these minimum and maximum values for the absorbed dose, an exposure strategy has to be developed. At the interface between resist and substrate, the lowest dose possible should be absorbed. Higher doses would lead to excessive production of secondary electrons generated at the interface by the X-rays. These secondary electrons break the chemical bonds between substrate and resist even in the neighbouring unexposed areas, leading to a loss of adhesion between substrate and resist. On the other hand, a dose as high as possible is preferred at the resist top surface for minimizing exposure times. This top-to-bottom ratio depends on two parameters only, the thickness of the PMMA resist and, via the absorption spectrum of PMMA, the wavelength of the X-rays. For a given syn-

8.2 The LIGA Process for Mold Inserts Fig. 8-8 SEM picture of a 2.5 mm high microstructure (PMMA, GS233)

chrotron source with its synchrotron radiation spectrum, high-energy photons can be cut off by mirrors at specific grazing angles in the beamline. Low-energy photons are absorbed in filters in the beamline or between beamline and the mask. Mirrors and filters define a bandpass and allow exposure of the resist to an optimized dose ratio. It is worth mentioning that in exceptional cases such as for the ANKA facility, the particle energy of the synchrotron can be freely adjusted for generating an optimum spectrum of the source. The number of such main user shifts for LIGA is, however, limited since special operation of the accelerator affects all other users at the facility. Table 8-2 shows the LIGA beamlines available at ANKA [24] and gives their characteristic wavelengths and the range of heights for which they have been designed. Although excellent results have demonstrated the extendibility of LIGA to very high structures (see Fig. 8-8), the processes for ultra-deep X-ray lithography are still at a development level [25, 26]. The DoseSim code mentioned above calculates the X-ray-spectrum from a bending magnet of a synchrotron. The modifications of this spectrum are calculated for inserting X-ray mirrors and filters according to the specific beamline. Also, the program is highly flexible and allows the user to calculate additional data, for example optimizing the absorber thickness for a desired dose after the absorber, filter options for a desired ratio top dose/bottom dose and expected time for development. The program also provides additional information such as the dose profile or the spectral distribution of the radiation after each absorber. 8.2.5 Development Process

After irradiation, the substrate with the exposed resist is placed in an organic developer solution, where the low molecular weight domains are being dissolved [27]. The development is completed by a rinse sequence to remove all residual organic components which starts in BDG (butyldiglycol) and ends in dis-

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8 Lithographic Fabrication of Mold Inserts Fig. 8-9 (a) Processing unit for wet-chemical development with megasonic agitation; (b and c) SEM pictures of details from microspectrometer structures (370 lm high) after development. (The lower picture demonstrates that the gluing zone behaves very similar to the resist in the lithography steps)

a

b

c

tilled water. The microstructures finally are dried in air in order to avoid any mechanical stress due to spinning [28]. The so-called GG developer goes back to an invention by Siemens AG [29] and is a mixture of 60% diethylene glycol monobutyl ether, 20% morpholine, 5% ethanolamine and 15% distilled water. The simple issue that all the exposed material must be removed turns out to be a critical point since components with fairly large molecular weight need to be removed with high certainty. Since the underlying mechanism for material transport is diffusion, the developing time depends strongly on resist thickness, exposure parameters, developer temperature, aspect ratio and layout of the structure and the type of bath agitation. The development process can be supported by application of megasound (Fig. 8-9 a). For a defect-free development of fragile microstructures, the megasound frequencies need to be in excess of the range where cavitation is observed. A typical frequency for megasonic agitation is 1 MHz. In some cases, also the intensity has to be adapted to the geometry of the microstructures [30]. Fortunately, the PMMA-GG resist developer system has an extremely high contrast so that the unexposed areas do not suffer even under severe overdevelopment [31]. This rather general statement will be commented on in Section 8.3.

8.2 The LIGA Process for Mold Inserts

8.2.6 Electroforming of the Mold Insert

Electroplating is the key step for the fabrication of metallic microcomponents and tools such as masks and molding tools. Established routine processes are often referred to by the kind of plating solution used. However, typically, every user performs the base processes with differences in electrolyte formulation and operation, for example due to specific fabrication environments such as the plating apparatus, to specific requirements on the mechanical properties or to the appearance of the surface. IMT uses mainly two fabrication sequences to obtain X-ray-lithographically patterned mold inserts, which are shown in Fig. 8-10. If the bottom of the substrate can be easily reached by the electrolyte, i.e. there are no tapped hole-like structures with high aspect ratio, a seed layer of 10 nm chromium and 50 nm gold is deposited all over the substrate by PVD techniques followed by the electroplating of the monolithic mold insert. For high aspect ratio nickel structures of the tapped hole type (columns on the mold), no seed layer is used but the structures are filled from the substrate [32] which carries the structured PMMA resist layer. At first, intermediate layers are plated into the resist cavities in the order nickel, gold, copper with thicknesses of 2, 1 and 5 lm, respectively. The plating processes are summarized in Table 8-3. The nickel–gold–copper layers fulfil functions which help to minimize defects, improve adhesion and optimize the surface quality of the mold insert. These different functions are summarized in Table 8-4 for the specified intermediate layers. The electroplating of the monolithic mold insert follows these steps. The first version, coating the whole resist template with a seed layer, leads to mold inserts of the highest quality in all details of the resist structure. Fig. 8-11

a

b

Fig. 8-10 Schematic depiction of the fabrication of the galvanic seed layer on the substrate. (a) Deposition of seed layers all over the substrate by PVD techniques; (b) deposition of intermediate layers into the resist cavities to guarantee good start conditions, excellent leveling and easy separation from the substrate

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8 Lithographic Fabrication of Mold Inserts Table 8-3 Purpose of intermediate layers used in standard LIGA mold insert fabrication

Layer

Thickness (lm)

Function

Nickel Gold

2 1

Copper

5

Homogenization of the substrate by increasing nuclei density Adhesion and protection. Copper can not be plated directly on to nickel with good adhesion Separation layer yielding very good surface finish

Table 8-4 Plating processes applied for fabrication of standard LIGA molding tools

Plating process

Current density (A/dm2)

pH

Temperature (8C)

Anode material

Nickel–sulfamate type Copper–sulfuric acid type Gold–sulfitic type

1–2 4

3.4–3.8 <1

52 21

S-depolarized Ni P-depolarized Cu

56

Pt-covered Ti

0.4

7

Fig. 8-11 SEM picture of a galvanic replicated grating (nickel, 370 lm high, width of grating teeth 5 lm, depth of grating teeth 0.2 lm)

shows this for a 370 lm high nickel structure with a 0.2 lm grating replicated from the resist sidewall. The limit in the aspect ratio for tapped hole or valley structures is due to the nickel growth rate being different for different directions. This will lead to voids or cavities at even moderate aspect ratios if the electroplating can start from a seed layer that covers all of the resist sidewalls and, in particular, the top surface. The second version of having only the substrate surface conductive and the resist surface insulating overcomes this limitation but it comes with a drawback which can be significant for applications: when the growth from the substrate reaches the top, the discrete nickel sections must somehow become connected. Since the metal growth is also sideways, this process can still be used to form a monolithic metal part.

8.2 The LIGA Process for Mold Inserts a

b

Fig. 8-12 SEM pictures of galvanic seams in structural cavity of a nickel mold insert

When the metal grows over the insulating surface, the fronts originating from different growth centers will finally meet and seams will be formed. Such seams will become filled with polymer during molding and upon demolding give rise to defects in the structures or, even worse, significantly raise the demolding forces so that neither a reliable nor a repeatable process window for demolding can be found. Unfortunately, the upper limit for the distance between adjacent growth centers is only 200–300 lm for which no seams appear. This maximum distance requirement introduces a severe limitation on layouts for LIGA mold inserts. Fig. 8-12 shows rather mild seams in an SEM picture. Ways and limitations for circumventing this effect will be discussed in Section 8.3. Nickel electroforming of mold inserts is performed using commercial production equipment [33] with an electrolyte volume of V = 40 L. The electrolyte used at IMT is a boron acid nickel sulfamate electrolyte (T = 52 8C, pH 3.5–3.7), the same as for other microelectroplating applications at IMT (for more information see [34] and Chapter 14). The electrolyte is agitated by constant circulation and filtration of the solution, while the substrate itself is moved backwards and forwards for convection. Originally, for mechanical stability, the overall thickness of the nickel body was chosen to be 5 mm. This results in plating times of up to 2–3 weeks depending on pattern geometry. During plating, the process has to be maintained within the process window with respect to the parameters pH value, surfactant concentration and concentration of degradation products. Further, the plating equipment must ensure constant temperature, convection and water replenishment resulting in a constant electrolyte volume during plating. During the growth inside the resist cavities, a critical current density must not be exceeded, because a rising pH value inside such cavities can result in nickel hydroxide formation, which inhibits further nickel deposition. This current density limit is *1 A/dm2, which corresponds to a growth rate of 12 lm/h. When the metal reaches the top of the resist layer, the current density should be increased to cover the additional area. A larger growth rate can also be applied, because the mass transfer is stronger and better controlled compared with the situation inside the cavities [35].

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8 Lithographic Fabrication of Mold Inserts

Fig. 8-13 Trend chart for monitoring of the electroplating unit. The horizontal lines span the warning range and the control limits. Deflection includes difference in thermal expansion coefficients of Si and Ni

The residual stress developed during nickel plating is below 50 MPa, which is low compared with conventional plating processes. Since this internal stress is a major source of deformation of the mold insert after substrate removal, the stress needs to be regularly measured in plating stations. For monitoring stress a method was developed where nickel is deposited on a metallized 4-inch Siwafer under standard conditions of 52 8C at regular time intervals (normally monthly). Stress in the nickel layer and the difference in thermal expansion of Si and Ni will deform the plated wafer after removing it from the bath and cooling it to room temperature. This deflection is determined and serves as a measure for inherent stress in the nickel film after correction for the thermal expansion effect. Based on the monitoring results, monitor charts for IMT’s plating stations including upper and lower limits were introduced (an example is shown in Fig. 8-13 [36, 37]). The application of stress-reducing agents would involve degradation of these species during plating, which would make the efforts much more tedious to control the plating baths by analytical methods in addition to the stress monitoring. Additives for leveling of the surface are not applied for the same reason. Since electroplating is a wet-chemical process at elevated temperatures, special care must be taken that areas which need to be protected are perfectly sealed off. A picture of a mold insert as it comes out of the plating station is shown in Fig. 8-14 together with the plating holder [38–40]. The holder needs to provide a reliable electric contact, needs to be ‘electrolyte proof’ over the whole duration of the process and it must allow for easy and reliable mounting and demounting of the resist–substrate and the substrate–resist–metal packages, respectively.

8.2 The LIGA Process for Mold Inserts a

b

c

Fig. 8-14 (a) Substrate holder for electroforming of mold inserts to facilitate loading and unloading of the substrate, there are two vacuum lips to seal off substrate and cover independently); (b) holder with substrate after nickel electroforming; (c) substrate after unloading from the holder (nickel height *6 mm)

8.2.7 Machining the Mold Insert to its Final Contour

After plating, the nickel piece is machined to the desired outer dimensions by electrodischarge machining (EDM) (see Fig. 8-15). IMT’s standard dimensions including tolerances are given in Fig. 8-16. Since the mold inserts will typically have to be precision mounted into the holder that fits a hot embossing or a injection molding machine, the required tolerances are sometimes challenging, for example, ±3 lm. Such tolerances can only be met by precision machining of the mold insert countour and the holder’s contour. After machining the nickel body is mechanically separated from the copper substrate. The separation occurs at the Ni–Au interface (see also Section 8.2.6). Because of that, the nickel surface of the mold insert is coated by the Au–Cu intermediate layer. These layers are removed by dissolving the underlying copper with a selective etch material, which is composed of NH3 and ClO2– [41]. Another possibility for separating the nickel body from the thick copper substrate is to cut the substrate by EDM. In this case, a residual copper layer (with a thickness between 200 and 500 lm) remains on the surface of the mold in-

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8 Lithographic Fabrication of Mold Inserts

a

b

Fig. 8-15 Schematic depiction of machining the outer dimensions of mold inserts by EDM

sert. This layer is etched (see above) followed by a second wet-chemical etch step in 5% HF solution to remove the titanium layer. 8.2.8 Polishing the Mold Insert

The next step in the manufacturing chain for molding tools is the polishing of the surface (see Fig. 8-17 a). This step can be omitted for mold inserts electroplated with intermediate layers. However, even in these cases, polishing will improve the surface quality and will level out possible imperfections related to intermediate layer inhomogeneities. For mold inserts with electroplating starting directly on the titanium oxide, polishing serves two purposes. First, the roughness of the surface which may complicate demolding because of mechanical interlocking between the polymeric molding material and microcavities at the mold surface is reduced (see Fig. 8-17 b). Second, although in many cases not relevant, the polishing will remove *10 lm of metal, which corresponds to the thickness of the resist glue. In the glue area, slightly rougher sidewalls and undercuts of 100 nm have been observed. For mechanical polishing, the cavities on the surface must be filled with a special casting resin, based on PMMA. Polishing is done with three different diamond suspensions (9, 3 and 1 lm) and three different polishing cloths at a polishing machine. With this process step a surface roughness of < 30 nm is obtained.

8.2 The LIGA Process for Mold Inserts

3.1±0.1 0.05 B 0.05 A

A

28–0.02

–0.1

30–0.2

Ra 1.6

Ra 1.6 B

0.05 B

` 0.04

0.05 A ±0.1

5

Fig. 8-16 IMT’s standard outer dimensions (including tolerances) for LIGA mold inserts

8.2.9 Cleaning of the Mold Insert

In the last manufacturing step, the PMMA resist material and/or the casting resin must be removed from the cavities of the mold insert. The process employed today is based on an X-ray flood exposure, followed by ultrasonic-supported stripping of the polymer in ethyl acetate [42]. At a temperature of 78 8C (boiling-point of ethyl acetate), the duration is *1 h. Finally, several rinsing steps are necessary to prevent any dissolved polymer from agglomerating on the mold insert. When contacted with air, it becomes virtually impossible to remove such residue.

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a

8 Lithographic Fabrication of Mold Inserts b

Fig. 8-17 Nickel mold inserts with two different microspectrometer layouts: (a) electroplated with intermediate layers, unpolished surface; (b) electroplated on PVD layers, polished surface (see section 8.2.8)

a

b

c

d

Fig. 8-18 Examples of residue on LIGA sidewalls after stripping and various cleaning procedures. (The origin is still not completely clarified.)

8.2 The LIGA Process for Mold Inserts

Fig. 8-19 Some examples of defects (particles, cracks, residues) in LIGA structures

The quality of the sidewalls make LIGA predominantly interesting for applications in which the sidewall is used for reflection (mirrors), diffraction (spectrometers) or refraction (lenses) of light. Reliable control of the sidewall quality, however, still poses a problem since scanning electron microscopy (SEM) is not suitable for in-line control, nor does it give quantitative information. Fig. 8-18 shows residues which cannot be removed if they show up. XPS and Auger spectra, which are not easy to take from such residue, are currently being analyzed in order to gain information on the possible origin of this contamination. Defects such as those shown in Fig. 8-19 are more typical. Residue as shown on the right can be significantly reduced by appropriate ultrasound treatment [43]. Naturally, cracks and particles, when incorporated into the metal structure, can no longer be removed. Such defects can be traced back throughout the process sequence and, therefore, their density needs to be monitored during all process steps.

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8.2.10 Mold Insert Specifications and Their Control

The mold insert must fulfil specifications in order to guarantee a desired functionality. In order to ensure the quality of the product, a function control must be executed but general guidelines for this process are not available. A perfect fabrication is virtually impossible. Therefore, the fabrication control at IMT includes an in-depth microscopic inspection of the layout area with a detailed floor plan of defects. SEM pictures are taken to give a general idea about what the defects look like (see Figs. 8-18 and 8-19). The planarity of the mold insert is verified after several key steps during the fabrication sequence whereas the metal hardness is measured only after the mold insert is finished. Critical issues are the dimensional parameters of the microstructures [44–46]. Their physical limitations are addressed in Section 8.3. The dimensions are measured using a Werth Coordinate Measuring Machine (CMM) [47], which allows for semiautomatic data acquisition. This equipment measures by image data processing and also by the Werth Fiber Probe (WFP). However, owing to various effects, for example the instability in illumination conditions, a calibration better than ± 2 lm against a national standard is still impossible [48]. Furthermore, if dimensional data from top to bottom are required, the distortion of the optics of the WFP becomes significant so that the data cannot be interpreted in a reliable way.

Table 8-5 IMT’s mold insert specifications

Feature

Set value (with tolerances)

Material Outer dimensions and height of the mold insert (after polishing) Hardness of material Structural depth

Nickel See Fig. 8-15

Minimal structure details (CAD) Lateral precision of microstructures

Maximum outer dimensions of layout area Height of mold insert Parallelism of structured front side to backside Flatness of structured front side (before and after polishing) Angle of structured front side to side face Roughness of structured front side (after polishing) Roughness of structures on bottom Surface irregularities of the structured areas (see DIN ISO 10110-7)

³ 230 HV 0.1 Between 200 and 750 lm (tolerances between –0/+40 and –0/+10 lm) 0.2 lm 0–50 lm (± 1 lm) 50–5000 lm (± 20 lm) 5000–50 000 lm (± 20 lm) 20 ´ 60 mm 5 mm (± 100 lm) £ 50 lm £ 40 lm 908, –0.58 Ra < 30 nm Ra < 100 nm 5/50 ´ 0.016 per cm2

8.3 Advanced Issues in LIGA Fabrication

Table 8-5 summarizes IMT’s mold insert specifications as they can be guaranteed in the standard process. Other specifications may be established but probably will require additional development work.

8.3 Advanced Issues in LIGA Fabrication 8.3.1 Precision-limiting Effects

In this section, the limits and the future development of LIGA mold inserts are discussed. First, the limit of precision is addressed. The original papers on LIGA [49] discussed the theoretical limits of X-ray-lithography. The conclusions were that Fresnel diffraction – increasing with increasing X-ray-wavelength and leading to dose deposition under absorber edges – and the creation of secondary photoelectrons – decreasing with increasing wavelength – are the two main effects limiting resolution. The minimum of the combined effect occurs for typical LIGA exposure conditions at wavelengths of 0.3 nm and numerically limits the resolution to only 100 nm. In fact, X-ray proximity lithography has been demonstrated to produce feature sizes far below 100 nm for very thin resist layers as they are used for semiconductor lithography. In LIGA, however, practical aspects have proven to impose much larger dimensional deviations or more complex effects on the resist profile than predicted by diffraction and photo electron emission. Thermal loading of the mask and resist–substrate sandwich during to the scanning movement in the intense X-ray beam plays an important role in defining the width of the area which is neither fully exposed nor fully unexposed [50]. A reduced temperature of the development bath ensures highest resist contrast [51], which has the effect that areas with unintended dose exposures are hardly attacked by the developer. The details of the structure profiles depend strongly on the specific set-up and the impinging radiation flux and become more critical for very high resists. Up to 1 mm resist height, the lateral loss (or sidewall inclination) is of the order of 0.1 lm per 100 lm. For a 3 mm resist height, the lateral difference between top and bottom can be as large as 10 lm. Thermal loading also plays an important role in the resolution limits of sidewall details. An example is optical gratings (see Fig. 8-1 a and b), where 0.2 lm steps with ideal sharp corners have to be patterned over the full resist height of nearly 400 lm. In such a case, the exposure conditions are rather delicate. Another practical aspect is related to secondary radiation. Photoelectrons generated at the resist–substrate interface may significantly reduce adhesion between substrate and resist [52]. Further X-ray fluorescence from the mask membrane can introduce significant dose deposition on the surface of the resist, including presumably unexposed areas. Fig. 8-20a shows rounding of the right-angle resist structures due to this fluorescence. Fig. 8-20b gives an example

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a

8 Lithographic Fabrication of Mold Inserts b

Fig. 8-20 (a) Example of influence of X-ray fluorescence. (b) Use of X-ray fluorescence for patterning

showing that such effects may be used for specific purposes. Here, after a primary patterning step, the top of the structures has been exposed to fluorescent radiation with less penetration depth than the original radiation. The top part of the structures has a reduced molecular weight which is associated with a reduced melting-point. A controlled temperature increase leads to melting of the top parts only and surface tension makes them spherical. With this method, lenses or arrays of lenses can be fabricated. In addition to lithography, the subsequent electroplating step also introduces limiting effects. The dominant one, particularly because electroplating of all but a few metals requires elevated temperatures, is the distortion of the PMMA by thermal expansion and also swelling of the resist in aqueous solutions [53–55]. This effect is strongly dependent on the details of the layout and can amount to 5 lm deviations at 350 lm structure height. With respect to mold inserts, this effect is critical since it can lead to negative nickel sidewall angles, which are highly detrimental for the demolding process. Although often discussed, no convincing remedies have been reported to date. Over the last few years, a strong LIGA activity has been established at Sandia National Laboratories, Livermore, CA, USA (see the next section). At this laboratory, S. K. Griffiths and his team have focused on modeling precision limiting factors in LIGA [55–57]. Meanwhile, the models treating individual issues have been combined into a simulation program predicting the process influence on the profile evolution [58].

8.3 Advanced Issues in LIGA Fabrication

8.3.2 Future Goals to Advance LIGA Mold Insert Fabrication

IMT’s efforts towards a standardized process sequence have already had a significant impact on the reliability and repeatability of LIGA mold insert fabrication. However, as another consequence the process is difficult to adapt to specific customer requests. Hence the needs for further development have been identified: 1. Resist top metallization for quasi-seamless mold inserts. 2. NiFe or NiCo as additional materials for mold inserts. Such molds will allow for replication at higher temperatures than is possible with Ni molds. 3. Increased pattern area with 70 mm diameter. With regard to reducing or eliminating seams, several attempts have been made in the past. The common idea is to deposit a patterned metal layer on top of the structures without coating the sidewalls of the resist pattern. One of the main problems in this approach is sufficient adhesion of the metal layer to PMMA. The metal layer has to function in the harsh environment of the electrolytic bath at least until the electroplated Ni reaches the top of the structures. Many molding materials require temperatures beyond 200 8C during replication. The degradation of the hardness for Nickel (HV0.1&300) sets in slightly above 200 8C. NiFe, on the other hand, has an initial hardness of HV0.1&550, which increases to 700 for temperatures up to 400 8C, thus bringing new applications within LIGA’s reach (see also Chapter 14). However, the corrosion stability of this alloy depends significantly on the content of saccharin which is introduced to reduce stress and therefore to obtain highly planar mold inserts [59, 60]. The initial hardness of the mold insert is often taken as a main indication for characterizing the lifetime. However, failure mechanisms have not yet been studied in sufficient detail. They may depend on many parameters, such as layout and load conditions. Some of the IMT mold inserts made from nickel have been used in an industrial manufacturing environment several tens of thousands of times [61]. Lifetime issues, therefore, do not need not be addressed with high priority. For a discussion of tribological aspects, see Chapter 21. Finally, increasing the area seems fairly straightforward at first glance. A major step has been taken recently by making beryllium masks available [62], which offer an area of nearly 80 mm diameter. However, many technical details need to be solved for transferring essentially all process steps from the primary (or intermediate) mask to resist deposition and all the other steps to a new format. Most important, in-depth experience is essential for acceptable process yields, which tend to decrease drastically with paradigm changes [63].

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8.3.3 Alternative to X-Ray LIGA: UV Lithography

With the advent of the UV-sensitive SU-8 resist family [64, 65], the LIGA concept has been extended to UV lithography, with UV radiation replacing X-rays but also followed by subsequent electroplating and molding. Metallic components [66–69], in addition to mold inserts [70, 71], can be fabricated by using conventional UV lithography and subsequent electroplating. SU-8 is a negative tone, epoxy-based, near-UV photoresist that was originally developed and patented by IBM. Owing to its low optical absorption in the UV range, this photoresist can be exposed in very thick films, even beyond 1 mm thickness. The SU-8 resist contains a few percent of photoacid generators that will produce a strong acid upon absorption of a photon. This photoacid acts as a catalyst for the cross-linking reaction that takes place during a post-exposure bake. The basic characteristics of this resist are: · it can be spun to a thickness of 1 mm and beyond; · it has been optimized for exposure at 365 nm wavelength where the standard high-pressure Hg lamp has a high-intensity peak; · it has an extremely high sensitivity so that relatively short exposure times are required; · it provides high resolution; · its processing may be integrated into standard MEMS process lines; · it has a high thermal stability and good chemical resistance. However, the major drawback is linked to the last advantage in that it is insoluble in most chemicals, which means it is difficult to remove [72]. Many solutions have been offered in the literature for SU-8 stripping but none is really

a

b

Fig. 8-21 (a) Detail of nickel mold insert for a fluidic application made by UV LIGA (structures 50 lm high and 30 lm width). (b) Detail of nickel mold insert for an optical application made by UV LIGA (structures 50 lm high and 50 lm width)

8.4 Conclusions

convincing. At IMT, an innovative approach was successfully demonstrated by using laser micromachining to remove thick layers of SU-8 on silicon wafers or mold inserts [73]. The drawbacks of this method are the limitation of the aspect ratio to *2, its processing time and its cost. The advantages for users of the EPON SU-8 technology are the use of a standard chrome–quartz mask and a large functional area when combined with a 6-inch wafer substrate. In contrast to X-ray-LIGA masks, chrome masks are readily available within less than 1 week through various suppliers, which is an additional advantage with respect to delivery times. The maximum diameter for a mold insert at IMT is 120 mm. The smallest feature sizes are 5 lm and the structure heights lie between 10 and 100 lm [70]. Roughness values for sidewalls, bottom and top of the mold insert are below 80 nm. Application examples are capillary fluidic devices, lab-on-chip applications and optical waveguides (see Fig. 8-21). The potential for fabricating twolevel structures with this process has also been also demonstrated [74].

8.4 Conclusions

Following its invention in the 1980s, the potential of LIGA became evident in the 1990s when numerous examples of structures and devices were presented. The lead in MEMS technologies, however, was assumed by silicon micromachining and not by LIGA, in spite of its technical superiority for many applications. The main reasons for this are the technical infrastructure and process know-how required and cost issues. Silicon-based technologies could exploit the vast technology base developed for chip-making with billions of dollars invested. LIGA, on the other hand, was new and confined to research laboratories. Furthermore, as a key process step, LIGA required access to a synchrotron radiation facility, again a research laboratory and often unacceptable for industries establishing manufacturing plants. Nevertheless, and this is a demonstration of the technical strength and superiority of LIGA, several industrial LIGA products have been launched by industries and synchrotron radiation facilities are currently widely used for microfabrication (see Table 8-6). As a consequence of the LIGA weaknesses – a long process sequence with highly specialized individual processes, need for synchrotron radiation sources, cost and throughput issues – the LIGA developers focused their efforts in past years on quality management, stabilization of processes, standards and cost. LIGA laboratories such as the IMT at Forschungszentrum Karlsruhe and IMM in Mainz introduced rigid quality management systems following ISO 9001 : 2000. Small volume production was demonstrated and reliability of processes was dramatically increased. The concept of ‘deliver mold inserts on demand with quality as specified’ was verified and small- and medium-sized companies increasingly accept that they can now simply purchase a mold insert for their production.

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8 Lithographic Fabrication of Mold Inserts Table 8-6 LIGA facilities worldwide

Country

Facility/provider

Synchrotron radiation facilities with LIGA activities Australia Australian Synchrotron – under construction Brazil LNLS Canada CLS – under construction France SOLEIL – under construction Germany ANKA BESSY-II Italy Elettra Japan NewSubaru, AURORA Korea PSL Russia VEPP-III Singapore SSLS Sweden MAX-II Taiwan NSRRC USA ALS APS CAMD NSLS-II SPEAR-III SRC Commercial LIGA providers Germany ANKA GmbH, Karlsruhe BESSY GmbH, Berlin USA Axsun Technologies Inc., Livermore, CA International Mezzo Technologies, Inc., Baton Rouge, LA

Further information

www.synchrotron.vic.gov.au www.lnls.br www.cls.usask.ca www.synchrotron-soleil.fr. www.fzk.de/anka www.bessy.de www.elettra.trieste.it www.spring8.or.jp www.ritsumei.ac.jp http://pal.postech.ac.kr http://ssrc.inp.nsk.su http://ssls.nus.edu.sg www.maxlab.lu.se www.srrc.gov.tw http://www-als.lbl.gov http://www.aps.anl.gov/ http://www.camd.lsu.edu http://www.nsls2.bnl.gov http://www2.slac.stanford.edu http://www.src.wisc.edu/ www.fzk.de/anka www.bessy.de www.axsun.com www.mezzotech.biz

Furthermore, new concepts are currently being tested where each component is produced directly by X-rays, similar to semiconductor manufacturing, at acceptable cost. This so called ‘direct LIGA’ approach uses only the first (LI) or the first two (LIG) steps of LIGA, thus reducing the complexity of mold fabrication and replication. The major concerns then are cost and throughput. Two highly encouraging approaches are pursued: new resists such as SU-8 reduce the exposure times by one to two orders of magnitude and the ability to expose very large areas in one exposure step. The latter concept has already led an industrial consortium to propose a mostly automated fabrication line at the ANKA facility at Forschungszentrum Karlsruhe. In addition to the described efforts to make LIGA acceptable as a manufacturing technology for a large variety of industrial products, cutting edge research in and with LIGA remains a hot topic. The goals include research in new mate-

8.6 References

rials, new replication techniques and new lithography approaches exploiting the short-wavelength nature of X-rays. Questions such as ‘how small can we really get’ and ‘can we overlap our top-down technologies with the typical bottom-up approach in nanotechnology?, and at which dimensions?’ are currently being addressed. Making devices much smaller than is possible today will open up entirely new fields in research and applications. Examples may include photonic crystals, X-ray microscopes and devices for research with individual molecules, cells and more.

8.5 Acknowledgment

LIGA technology has been developed at IMT for more than a decade. It has been a truly collaborative and interdisciplinary effort by scientists, engineers, technicians and support personnel. Without their vision in LIGA as a widely used technology for microsystems fabrication, IMT would not have been able to keep its technological leadership in this field. We are, therefore, deeply indebted to IMT’s current and former co-workers.

8.6 References 1

2

3

4

5

6 7

8

E. W. Becker, W. Ehrfeld, P. Hagmann, A. Maner, D. Münchmeyer, Microelectron. Eng. 1986, 4, 35–36. H. Guckel, K. Fischer, E. Stiers, B. Chaudhuri, S. McNamara, M. Ramotowski, E. D. Johnson, C. Kirk, Microsyst. Technol. 2000, 6, 103–105. J. Mohr, A. Last, U. Hollenbach, T. Oka, U. Wallrabe, J. Lightwave Technol. 2003, 21, 643–647. T. Hanemann, M. Ade, M. Börner, G. Motz, M. Schulz, J. Hausselt, Adv. Eng. Mater. 2002, 4, 869–873. O. V. Makarova, D. C. Mancini, N. Moldovan, R. Divan, C.-M. Tang, D. G. Ryding, R. H. Lee, Sens. Actuators A 2003, 103, 182–186. M. Heckele, W. K. Schomburg, J. Micromech. Microeng. 2004, 14, R1–R14. F. J. Pantenburg, S. Achenbach, J. Mohr, J. Vac. Sci. Technol. B 1998, 16, 3547– 3551. P. Krippner, T. Kühner, J. Mohr, V. Saile, Proc. SPIE 2000, 3912, 141–149.

9 C. Müller, O. Fromheim, J. Göttert, T.

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23 24

25 26 27 28 29 30 31 32

33 34 35

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www.roehm.de/de/plexiglas/index.html German Patent DE 4 230 297, 1993. P. Meyer, C. Cremers, A. El-Kholi, D. Haller, J. Schulz, L. Hahn, S. Megtert, Microsyst. Technol. 2002, 9, 104–108. P. Meyer, L. Hahn, J. Schulz, Rev. Sci. Instrum. 2003, 74, 1113–1119. For further information about ANKA and useable beamlines, see http:// hikwww1.fzk.de/anka/ANKA_BB.pdf S. Achenbach, T. Mappes, J. Mohr, J. Vac. Sci. Technol. B, 2004, 22, 3196–3201. S. Achenbach, Microsyst. Technol. 2004, 10, 493–497. A. El-Kholi, J. Goettert, J. Mohr, P. Bley, Microelectron. Eng. 1993, 21, 271–274. P. Meyer, A. El-Kholi, J. Schulz, Microelectron. Eng. 2002, 63, 319–328. US Patent 4 393 129, 1983; German Patent DE 3039110C2, 1990. A. El-Kholi, J. Mohr, R. Stransky, Microelectron. Eng. 1994, 23, 219–222. S. Achenbach, University of Saskatchewan, Canada, personal communication. W. Bacher, K. Bade, B. Matthis, M. Saumer, R. Schwarz, Microsyst. Technol. 1998, 4, 117–119. For further information: HEGA Apparatebau GmbH, Straubenhardt, Germany. K. Bade, Galvanotechnik 1999, 90, 801– 809. K. Bade, K. Leyendecker, A. Thomes, W. Bacher, in: Magnetic Materials, Processes and Devices IV, L. T. Romankiw, D. A. Herman (eds); Pennington: Electrochemical Society, 1996, pp. 697–708. K. Bade, U. Köhler, B. Krevet, B. Matthis, J. Schulz, presented at the EASTForum, Lyngby, Denmark, 10–12 April 2003. J. Schulz, K. Bade, M. Guttmann, L. Hahn, A. Janssen, U. Köhler, P. Meyer, F. Winkler, Microsyst. Technol. 2004, 10, 419–422. For further information, see www.silicet.de M. Guttmann, H. Moritz, Galvanotechnik 2002, 93, 2398–2402. M. Guttmann, K. Bade, B. Matthis, J. Schulz, H. Moritz, in: Proceedings of Oberflächentage 2004, 42. Jahrestagung der DGO 2004, Dresden; 2004, pp. 301–306.

41 A. F. Bogenschütz, U. Büttner, J. L.

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43 44 45 46

47 48

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Jostan, A. Marten, Galvanotechnik 1978, 69, 690–701. M. Guttmann, F. Winkler, J. Schulz, G. Hielscher, Metalloberfläche 2005, in preparation. For further information, see www.hielscher.com B. L. Dearth, R. G. Steinhoff Jr, Microsyst. Technol. 2003, 9, 197–203. J. F. Nichols, T. R. Kurfess, Microsyst. Technol. 2004, 10, 556–559. R. Thelen, P. Meyer, J. Schulz, P. Jakobs, J. Lange, M. Arendt, Wiss. Ber. FZKA 2004, 6990, 185–186. For further information, see www.werthmesstechnik.de R. Thelen, P. Meyer, A. Janssen, J. Schulz, EUSPEN2005 Book of Abstracts, ID177. D. Münchmeyer, PhD Thesis, Universität Karlsruhe; Wiss. Ber. Kernforschungszentrum Karlsruhe 1984, KfK 3732. S. Achenbach, PhD Thesis, Universität Karlsruhe. Wiss. Ber. FZKA 2000, 6576. F. J. Pantenburg, S. Achenbach, J. Mohr, J. Vac. Sci. Technnol. B 1998, 16, 3547– 3551. F. J. Pantenburg, J. Chlebek, A. El-Kholi, H. L. Huber, J. Mohr, H. K. Oertel, J. Schulz, Microelectron. Eng. 1994, 23, 223–226. A. C. M. Ruzzu, PhD Thesis, Universität Karlsruhe; Wiss. Ber. FZKA 2000, 6514. A. Ruzzu, B. Matthis, Microsyst. Technol. 2002, 8, 116–119. S. K. Griffiths, J. A. W. Crowell, B. L. Kistler, A. S. Dryden. J. Microeng. Micromech. 2004, 14, 1548–1557. R. H. Nilson, S. K. Griffiths, J. Electrochem. Soc., 2002, 149, G286–G296. S. K. Griffiths, J. Micromech. Microeng. 2004, 14, 999–1011. F. J. Pantenburg, Forschungszentrum Karlsruhe, personal communication. M. Guttmann, K. Bade, in: Jahrbuch Oberflächentechnik, A. Zielonka (ed.); Isernhagen: Giesel-Verlag, 2001, Band 57, pp. 103–120. Y.-M. Yeh, G. C. Tu, T.-H. Fang, J. Alloys Compd. 2004, 372, 224–230. STEAG microParts GmbH, personal communication.

8.6 References 62 H. Hein, A. Janssen, B. Matthis, P. Ja-

63 64 65

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kobs, S. Achenbach, D. Maas, P. Henzi, P. Meyer, ’Wiss. Ber. FZKA 2004, 6990, 183–184. W. Hilberg, Phys. Unserer Zeit, 1986, 17, 18–28. C. G. Khan Malek, Microelectron. J., 2002, 33, 101–105. J. Liu, B. Cai, J. Zhu, G. Ding, X. Zhao, C. Yang, D. Chen, Microsyst. Technol., 2004, 10, 265–268. H. Lorenz, M. Despont, N. Fahrni, N. LaBianca, P. Renaud, J. Micromech. Microeng., 1997, 7, 121–124. I. Song, P. K. Ajmera, J. Micromech. Microeng., 2003, 13, 816–821. T. Kohlmeier, V. Seidemann, S. Büttgenbach, H. H. Gatzen, Microsyst. Technol., 2002, 8, 304–307.

69 S. H. Oh, K.-C. Lee, J. Chun, M. Kim,

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S. S. Lee, J. Micromech. Microeng., 2001, 11, 221–225. S. J. Chung, H. Hein, Wiss. Ber. FZKA 2000, 6423, 195–196. R. Turner, S. Desta, K. Kelly, J. Zhang, E. Geiger, S. Cortez, D. C. Mancini, J. Micromech. Microeng. 2003, 13, 367–372. P. M. Dentinger, W. M. Clift, S. H. Goods, Microelectron. Eng. 2002, 61–62, 993– 1000. M. K. Ghantasala, J. P. Hayes, E. C. Harvey, D. K. Sood, J. Micromech. Microeng. 2001, 11, 1233–1239. T. Hirata, S. J. Chung, H. Hein, T. Akashi, J. Mohr, Proc. SPIE 1999, 3875, 164– 171.

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9 Material States and Surface Conditioning for Mold Inserts Ch. Horsch, V. Schulze, D. Löhe, Institute of Material Science and Engineering I, University of Karlsruhe (TH), Germany Abstract Micromilled mold inserts generally show roughnesses and surface textures due to the milling process. Micromilling in hardened steels generally leads to burr formation, which has to be avoided in mold and die manufacturing. The influence of heat treatment and machining parameters on roughness and burr formation is investigated to optimize the workpiece properties obtained. A further improvement of surface states and burr removal can be obtained by surface conditioning processes. An overview of such processes is given and the influence of peening processes on surface states and burrs is investigated. Keywords micromilling; machining properties; heat treatment; surface states; roughness; surface conditioning; micropeening

9.1 9.2 9.2.1 9.2.2 9.2.3 9.2.4 9.2.5 9.3 9.3.1 9.3.2 9.3.3 9.3.4 9.3.5 9.4 9.4.1

Introduction 222 Micro-side Milling of a Tool Steel

223

Specimens and Experimental Setup 223 Built-up Edges 225 Minimum Chip Thickness 227 Form Errors and Burrs 229 Modifications of the Microstructure 231 Micro-end Milling of Tool Steels

233

Specimens and Experimental Setup 233 Machining of Grooves 234 Roughness Analysis 235 Tool Wear and Burrs 237 Influence of the Process on the Material State Close to the Surface 238 Surface Conditioning of Micromilled Structures in a Tool Steel

Process Overview 239

Advanced Micro and Nanosystems Vol. 3. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31208-0

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9.4.2 9.4.2.1 9.4.2.2 9.4.2.3 9.4.3 9.4.3.1 9.4.3.2 9.5 9.6

Micropeening 241 Specimens and Experimental Setup Deburring 242 Surface States 243 Ultrasonic Wet Peening 244 Specimens and Experimental Setup Experimental Results 246

241

245

Conclusions 248 References 249

9.1 Introduction

Micromilling offers various advantages in manufacture of micromolds and dies. In contrast to lithographic processes that are mainly used by the semiconductor industry, it allows the manufacture of steel microstructures and hence the use of materials with high strength, toughness and wear resistance. Furthermore, milling allows the production of real 3D structures. As the use of diamond tools is hardly possible for micromilling of steels, micromilled structures made of steel show higher roughnesses than other milled microstructures. As carbide tools that are mainly used for milling of steels show a larger cutting edge radius than diamond tools, the formation of large burrs can occur in microcutting of steels. Especially in mold and die production, smooth surfaces are required and burrs are generally not tolerable. Therefore, surface conditioning after structuring is of high importance. In the first part of this chapter, the influence of the mechanical properties of the workpiece material on the surface state after micro-side milling is investigated. This process was chosen for basic investigations on micromilling mainly because of its simple kinematics and the large variability of machining parameters. As only straight grooves can be obtained by this kind of process, it cannot be used for mold manufacturing. In the second part, micro-end milling of different tool steels is investigated, with respect to roughness, tool wear, burr formation and modifications of the microstructure occurring in the surface layer. As end milling is generally used in micromold manufacture because it allows for complex geometries, the aim of the investigations on micro-end milling was the identification of material states showing good machinability. Further investigations on micro-end milling can be found in Chapter 5. In the third part, an overview of conditioning processes allowing roughness reductions and deburring is given and the possibilities for optimizing milled microstructures by dry and wet micropeening processes are shown.

9.2 Micro-side Milling of a Tool Steel

9.2 Micro-side Milling of a Tool Steel

For the basic investigations on surface generation by micromilling, a side milling process with a single-tooth tool was used, mainly for three reasons. First, the simple kinematics of the side milling process allow an easy comparison between the theoretical (kinematic) and the measured roughness values. Second, in contrast to multi-tooth tools, there can be no concentricity error between the different cutting edges. Third, this tool concept allows large tool diameters and hence a great variation of the cutting speed. 9.2.1 Specimens and Experimental Setup

The workpiece material investigated was the electro-slag remelted cold working tool steel 90MnCrV8 (German grade 1.2842), obtained as warm-rolled billets of 50 ´ 60 ´ 4000 mm. The milling specimens shown in Fig. 9-1 with dimensions 30 ´ 35 ´ 10 mm were manufactured in such a way that the feed direction in micromilling was perpendicular to the rolling direction. The material allowance for grinding was 0.1 mm on both the upper and lower sides. One quenched and two quenched and tempered material states were investigated. The heat treatment parameters and resulting hardnesses can be seen in Table 9-1. Micrographs of the different material states are shown in Fig. 9-2. The microstructure shows a high regularity and fineness, hence the material seems to be adequate for micromilling. In the micrographs, precipitations of M3C (M = Fe, Mn, Cr, V) that were not dissolved because of the low austenitizing temperature can be seen. The hardened specimen (H) and that tempered at 300 8C (T300) have a martensitic microstructure; that tempered at 600 8C (T600) shows a microstructure composed of ferrite and carbides.

Fig. 9-1 Geometry of the milling specimens

223

224

9 Material States and Surface Conditioning for Mold Inserts Table 9-1 Heat treatments and hardnesses of the materials investigated

Material

Notation

Heat treatment

Hardness (HRC)

90MnCrV 8

H T300

810 8C vacuum/1 h, ? oil 810 8C vacuum/1 h, ? oil, 20 min 300 8C 810 8C vacuum/1 h, ? oil, 20 min 600 8C

64 58

T600

36

Fig. 9-2 Microstructure of the investigated material states: quenched (left), tempered at 300 8C (middle) and tempered at 600 8C (right)

The mechanical properties of the three material states were determined by tensile testing and compressive testing. The fracture toughness was measured using CT specimens. Table 9-2 shows the mean values of the mechanical properties. Owing to the brittle character of the state H, the measured ultimate tensile strength is lower than the values expected from the hardness. Therefore, for comparisons of the mechanical strength, the compressive proof stress is used. The milling experiments were conducted at the Institute of Production Science (wbk) on a micromill [1] with a maximum revolution speed of 6000 min–1 and a maximum spindle power of 3.8 kW. The tools used for machining the grooves were manufactured of rhombic cemented carbide inserts by grinding. Their geometry can be seen in Fig. 9-3.

Table 9-2 Mechanical properties

Property

H

T300

T600

Rm (MPa) Rp 0.2 (MPa) A (%) Rdp 5 (MPa) Rdp 10 (MPa) KIc (MPa m½)

(1273) – – 4000 4300 36

2133 1907 2.2 3450 3650 47

1065 997 9.7 1150 1280 155

9.2 Micro-side Milling of a Tool Steel

Fig. 9-3 Micromilling inserts

They were installed with an inclination of 58 in order to have an effective rake angle of 88 and a clearance angle of 38. For measuring the microhardness of surface layers, a Fischerscope H100 hardness tester with a Vickers indenter at a maximum load of 100 mN was used in order to measure force-indentation courses. The Martens Hardness [2] is calculated by dividing the maximum test force F by the surface area of the indenter during penetration beyond the original surface of the test piece. The hardness values obtained were converted into Vickers hardness. 9.2.2 Built-up Edges

In machining processes with low cutting speeds and soft materials, local welding between the workpiece material and the clearance surface, so-called built-up edges, can occur. This results in a fluctuating tool geometry, tearing of workpiece material out of the generated surface and local material deposition on the generated surface, all leading to deteriorated surface states and reduced tool lifetimes. Generally, the extent of built-up edges decreases with increasing cutting speed because the diffusion processes between the tool and the workpiece leading to this effect are hindered owing to the lower contact times. Fig. 9-4 shows the topography of micromilled grooves in the three heat treatment states investigated at different cutting speeds. As can clearly be seen, the surface quality increases with increasing cutting speed and increasing workpiece hardness. Fig. 9-5 shows the measured roughness Rz [3] as a function of the cutting speed for the three material states investigated. In the quenched material state, initially the roughness decreases slightly with increasing cutting speed. Built-up edges can be considered as non-significant for the roughness at cutting speeds higher than 50 m/min. This result copes with the surface topography which is mainly influenced by striations due to the tool grinding. At cutting speeds above 350 m/min, the roughness increases again. Metallographic sections (cf. 9.2.5) prove that this roughness increase is

225

226

9 Material States and Surface Conditioning for Mold Inserts State

vc = 5 m/min

vc = 140 m/min

vc = 260 m/min

vc = 420 m/min

H

T300

T600

Fig. 9-4 Topography of microgrooves milled at different cutting speeds and with different material states

Fig. 9-5 Roughness Rz vs. cutting speed

due to thermal softening coupled with a stronger tendency to built-up edges. Built-up edges at a cutting speed of 420 m/min could be detected by EDX analysis of a cutting edge after use. The presence of iron in the cemented carbide tool proves that diffusion takes place between workpiece and tool.

9.2 Micro-side Milling of a Tool Steel

In the state T300, the roughness decreases first for increasing cutting speeds above 60 m/min. It increases again at cutting speeds above 250 m/min. Hence this softer material shows a stronger tendency for built-up edges. In total the roughness values are increased compared with the quenched state. In the state T600, high roughness values up to more than 7 lm are obtained for low cutting speeds. The roughness decreases with increasing cutting speed, but even for highest speeds the surface remains deteriorated by built-up edges. Only at vc = 420 m/min could surface qualities similar to those at the other material states be observed. Generally, the best surface qualities can be obtained on materials machined with high cutting speeds. 9.2.3 Minimum Chip Thickness

The kinematic roughness in side milling, resulting from the cycloid-shaped relative movements between tool and workpiece usually is described (e.g. [4]) by Rkin ˆ

fz2 8r

…1†

with the feed per edge fz and the tool radius r. For high chip thicknesses, this equation complies with measured roughnesses (e.g. [5]). For small chip thickness, the measured roughness can exceed the kinematic values by a factor of six [6]. The minimum chip thickness is assumed to be one of the main reasons for this phenomenon. First investigations on the effects of the minimum chip thickness were carried out by Sokolowski [7], Pahlitzsch and Semmler [8] and Dornhöfer [9]. In up-milling, at the entry of the cutting edge into the workpiece, at a cutting depth of zero, there is an elasto-plastic deformation of the workpiece before the cutting itself begins. Chip formation occurs only beyond a critical value called the minimum chip thickness, hmin. A springback of the elastically deformed surface is assumed to increase the roughness. In down-milling, i.e. milling with a decreasing chip thickness, the cutting process is interrupted at the minimum chip thickness and there is an elasto-plastic deformation of the workpiece surface. Fig. 9-6 shows the geometric relations between the minimum chip thickness and the surface roughness. From these geometric considerations, the following equation was found:   fz2 hmin rhmin Rth ˆ ‡ 1‡ 2 8r 2 fz

…2†

The first term represents the kinematic part of the roughness, which would tend to zero for very small feeds or high tool diameters. The second term introducing the minimum chip thickness counteracts this and becomes determinative for the roughness at small feed rates.

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9 Material States and Surface Conditioning for Mold Inserts

Fig. 9-6 Geometric relations between the minimum chip thickness and the theoretical roughness

The minimum chip thickness has been found to decrease with the cutting edge radius [4, 7, 9, 10]. Contrary to first assumptions that the minimum chip thickness equals the cutting edge radius, e.g. [11], most investigations showed that it is possible to have a cutting process at much smaller chip thicknesses [12, 13]. A first approach [14] is a comparison between the cutting process with a rounded cutting edge with a radius q and the rolling of steel. From equilibrium conditions shown in Fig. 9-7, it is assumed that a chip is formed if the resulting force has a component in the positive y direction (out of the workpiece surface). The extreme value for which the normal forces on the tool are equal is found at h = p/4, which results in a minimum chip thickness given by

Fig. 9-7 Equilibrium conditions on the cutting edge at the minimum chip thickness

9.2 Micro-side Milling of a Tool Steel Table 9-3 Roughness values and corresponding minimum chip thicknesses

State

Rz (lm)

hmin (lm)

H T300 T600

0.92 1.1 1.6

0.68 0.77 0.98

hmin ˆ r…1

cos h† ˆ 0:293r

…3†

As the geometric proportions at the cutting edge depend closely on the workpiece deformation and hence on its elasto-plastic properties, Equation (3) can only be an approximation. According to the considerations of Brammertz [4], the minimum chip thickness should decrease with the material’s ductility, so that a brittle material would show lower roughness values. For estimating the minimum chip thickness from the measured roughness values by Equation (2), it is necessary to eliminate the effects on the roughness due to built-up edges. In the hardened state, the roughness values between 100 and 300 m/min are nearly constant and their mean value can be considered as almost independent of built-up egdes. In the state T300, the roughness values show a minimum at a cutting speed of 260 m/min and the higher roughness values seem to be affected by built-up edges. In the state T600, the lowest roughness is measured for a cutting speed of 420 m/s, which will be used for estimating the minimum chip thickness. Table 9-3 gives the roughness values that have been used for calculating the minimum chip thickness for the three thermal treatment states and the hmin values calculated from these roughness values. Fig. 9-8 shows the estimated minimum chip thicknesses vs. the workpieces’ elongation to fracture, 5% compressive proof stress, hardness and fracture toughness. Nearly linear relations can be found between the minimum chip thickness and the mechanical workpiece properties. As these values are measures of the ductility and the resistance against plastic deformation, respectively, the hypothesis that the minimum chip thickness is mainly determined by the plastic properties of the material is confirmed. 9.2.4 Form Errors and Burrs

For the investigations of form errors and burrs, metallographic sections perpendicular to the feed direction were made. The micrographs in Fig. 9-9 show the form of the grooves which were intended to be rectangular with a depth of 150 lm and a width of 300 lm.

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9 Material States and Surface Conditioning for Mold Inserts

Fig. 9-8 Minimum chip thickness vs. mechanical properties

State

vc = 5 m/min

vc = 140 m/min

vc = 260 m/min

H

T300

T600

Fig. 9-9 Metallographic sections of grooves for different cutting speeds in the states H (top), T300 (middle) and T600 (bottom)

vc = 420 m/min

9.2 Micro-side Milling of a Tool Steel

In the quenched state, cracks could be detected up to cutting speeds of 260 m/min. These cracks were initiated close to the inner edge of the groove and propagated in an inclined direction towards the upper surface. The crack length decreases with increasing cutting speed. In the other material states investigated, no cracks could be found. In the quenched state, a pronounced rounding of the inner edges can be observed, which is due to stronger tool wear in this material state. In the two tempered states, only a weak edge rounding can be seen. At the quenched material, relatively large burrs were found whose base broadness decreased with increasing cutting speed. In the two other states, only little burr formation occurred. A clear relation with the cutting speed could not be found. Both tempered material states show good contour accuracy. The quenched state is much less appropriate for micromilling because of the higher tendency for burr formation and cracks. 9.2.5 Modifications of the Microstructure

For investigations on modifications of the microstructure effected by the milling process, the etched metallographic sections were examined with an optical microscope. In the hardened state, significant changes of the microstructure can be observed in Fig. 9-10. Close to the surface, a pale layer is found. Its thickness increases with the cutting speed. The layer below becomes darker owing to the etching procedure. Under these two layers, the material remains unaffected.

Fig. 9-10 Modifications of the microstructure in the hardened material state (vc = 340 m/min)

231

232

9 Material States and Surface Conditioning for Mold Inserts Table 9-4 Hardness distribution in the affected material

Layer

Hardness (HV)

Pale layer Dark layer Base material

1534 1034 1364

The inner groove edge, around which these layers are located, is machined by the tool corner between the minor and the major cutting edges. Since the highest temperatures develop there, one can assume thermal effects to be the origin of the microstructural modifications. To verify this hypothesis, the microhardness of the different layers was measured, yielding the results given in Table 9-4. The hardness distribution observed can be explained by the thermal effects during the milling process. The edge layer is heated and austenitized. Subsequently, it is self-quenched by the colder base material and thus retransformed into martensite. The thermal effect on the dark etching layer is lower, so that austenitization does not occur. The heating of this layer causes tempering of the material and hence a hardness decrease. In the base material, no tempering processes were detected. Because of the heating, which increases with the cutting speed, the thickness of the affected layers also increases with the cutting speed. The local heating also explains the decreasing tendency for cracking with increasing cutting speed. At low cutting speeds and hence at low material temperatures, the local stress induced by the cutting forces is released by unstable

Fig. 9-11 Burr formation from the austenitized edge layer (vc = 420 m/min)

9.3 Micro-end Milling of Tool Steels

crack growth affecting the whole area around the groove. With increasing cutting speed and temperature, the material strength and brittleness diminish and fracture toughness increases. Hence increasing cutting speeds favor plastic deformation instead of cracking. At the highest cutting speeds, the entire surface is austenitized and the very low strength of the austenite confines the plastic deformation mostly on this surface layer, thus leading to very narrow burrs (cf. Fig. 9-11). No similar microstructural modifications close to the surface could be found in the tempered material states. This is due to their higher fracture toughness and ductility, which inhibit cracking, and to their lower strength, leading to a lower deformation energy and thus to lower material temperatures.

9.3 Micro-end Milling of Tool Steels 9.3.1 Specimens and Experimental Setup

The specimens for micromilling were cuboids of 68 ´ 32 ´ 10.4 mm, which were obtained by sawing and milling of warm rolled plates of the vacuum-remelted hot-work tool steel X38CrMoV5-1 VMR (German grade 1.2343) and of Toolox 44®, a low-alloyed tool steel (30CrMo6 with 0.9% Mn and 0.6% Si) used for molds for polymers, which was delivered in a quenched and tempered state. X38CrMoV5-1 was used as it is one of the most common hot-work tool steels in precision mold and die manufacturing because of its good machinability and its low residual stresses and residual austenite content after tempering. Especially the high purity of the VMR grade is necessary for microcutting. Toolox 44 was chosen as a comparison with the standard material for machining and tribological properties. The X38CrMoV5-1 specimens were quenched and tempered or austempered. For quenching, they were pre-heated for 10 min at 850 8C, austenitized for 40 min at 990 8C in a vacuum furnace and quenched in oil. The hardness after quenching was 54 HRC. Subsequently, they were tempered twice at 520 or 560 8C for 120 min each under vacuum and cooled at 5 K/min in an air jet. Tempering at 520 8C is the manufacturer’s specification for work hardness. For austempering, the specimens were austenitized in a fluidized Al2O3 nitrogen bed at 990 8C for 30 min, quenched to 400 8C in an Al2O3 air bed of 20 8C and isothermally transformed in another fluidized Al2O3 air bed at 350 8C for 3 h. The hardnesses after the different heat treatments are listed in Table 9-5. After heat treatment, 0.2 mm was ground off from the upper and lower sides to remove distortions due to the heat treatments. Fig. 9-12 shows the microstructure of the investigated steels. As the microstructures of both quenched and tempered states of X38CrMoV5-1 are very similar, T560 is shown exemplarily.

233

234

9 Material States and Surface Conditioning for Mold Inserts Table 9-5 Heat treatments and resulting hardnesses

Heat treatment

Notation

Hardness (HRC)

Quenched and tempered at 520 8C Quenched and tempered at 560 8C Austempered to bainite Toolox 44

T520 T560 B T44

52 49 48 45

Fig. 9-12 Microstructure of the investigated steels: quenched and tempered (left) and austempered X38CrMoV5-1 (middle) and Toolox 44 (right)

Table 9-6 Material states and mechanical properties

State

Rp

T520 T560 B T44

1427 1375 818 1260

0.2

(MPa)

Rm (MPa)

A (%)

1698 1620 1768 1376

13.7 13.0 12.2 12.5

For tensile testing, specimens were produced by sawing and turning in a way that the specimen axis complied with the rolling direction. The mechanical properties (0.2% proof stress, ultimate tensile stress and elongation to fracture) of the investigated materials are given in Table 9-6. According to the similar hardness values, the four states investigated show similar tensile strengths and elongations to fracture. The austempered state shows a much lower yield strength and hence a high work hardening. 9.3.2 Machining of Grooves

For micromilling, a two-edged TiAlN-coated cemented carbide tool with a diameter of 300 lm, shown in Fig. 9-13, was used. The spindle revolution speed was 30 000 min–1 and the resulting cutting speed on the feed side was 57 m/min. A feed rate of 1 lm per tooth and a cut-

9.3 Micro-end Milling of Tool Steels

235

Fig. 9-13 Micromilling cutter

Fig. 9-14 Specimen with micromilled grooves

ting depth of 25 lm were used. Meander-shaped grooves with a total length of 550 mm as shown in Fig. 9-14 were machined by end milling. The dark color of the specimen is due to the ultrasonic cleaning process after milling that is necessary to avoid corrosion. 9.3.3 Roughness Analysis

For roughness analysis, the groove grounds were investigated with a confocal white-light microscope (Nanofocus lSurf). On every groove, 30 profile lines with a length of 320 lm were separated into roughness and waviness with a Gauss filter at a cut-off wavelength of 80 lm. Rz and Wt were calculated according to ISO 4287 [3] and a mean value over the 30 profile lines was calculated. The results are shown in Table 9-7.

Please check the numbers of cross-references throughout the text< /QA>

236

9 Material States and Surface Conditioning for Mold Inserts Table 9-7 Influence of the heat treatment state on roughness, waviness and profile depth

State

Rz (lm)

Wt (lm)

Pt (lm)

T520 T560 B T44

0.45 0.62 0.36 0.44

0.42 0.56 0.38 0.14

0.83 0.94 0.65 0.64

As can be seen, the softer state T560 leads to a higher roughness than T520. Although showing a similar hardness as T560, the bainitic specimen offers the lowest roughness of all heat treatment states investigated. T44 shows a roughness comparable to T520, but a very low waviness, hence the profile depth is nearly the same as for state B. The grooves’ ground (Fig. 9-15) shows nearly circular-shaped feed marks which are typical for end milling. In all states of X38CrMoV5-1, striae inclined to the feed direction can be seen, in addition to the feed marks. They are considered to be due to slight regenerative chatter, which is typical for surfaces milled with small feed rates [15]. The small cutter diameter leads to a low tool stiffness. Hence bending vibrations of the tool occur and the tool inclination due to the vibrations generates striae on the surface. These striae lead to an increase in the waviness values. Another indication for tool vibrations is the distance between the feed marks, which should be equal to the feed per edge or to twice this value in the case of a larger radial run-out. All investigated grooves show feed mark distances of about 7–10 times the feed per edge, which can only be explained by tool vibrations. The groove in Toolox 44 (Fig. 9-14, right)

0

50

100 150 200

250 300 lm

lm

0 50 100 150 200 250 300 350 400 450 500 550 lm

7 6.5 6 5.5 5 4.5 4 3.5 3 2.5 2 1.5 1 0.5 0

0 0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4 0.45 0.5 0.55 0.6 0.65 0.7 0.75

50

100

150 lm

mm

Fig. 9-15 Surface topography of micromilled grooves in X38 state B and Toolox 44

lm 3.2 3 2.8 2.6 2.4 2.2 2 1.8 1.6 1.4 1.2 1 0.8 0.6 0.4 0.2 0

9.3 Micro-end Milling of Tool Steels

237

shows a similar disposition of the feed marks with a distance greater than the feed per edge. Striae inclined to the feed direction cannot be found. From this it can be concluded that chatter phenomena are also present, but with a smaller magnitude, which can be explained by the lower strength of this steel. Despite these vibrations, the surface qualities obtained with profile depths smaller than 1 lm can be evaluated as very good for milled surfaces. This holds especially for state B of X38CrMoV5-1 and Toolox 44, which show the lowest Pt values. 9.3.4 Tool Wear and Burrs

Fig. 9-16 shows micrographs perpendicular to the grooves at different milled distances for the four material states investigated. Form errors and burr formations can be observed in these micrographs. As in end milling, tool wear mostly appears at the tool tip and at the minor cutting edges, leading to a rounding of the groove ground. In contrast, burr formation indicates wear of the major cutting edges. The high hardness of the specimen tempered at 520 8C leads to substantial tool wear, especially of the tool tip and of the minor cutting edges, which leads to a cutting depth that decreases strongly with the milling distance. With increasing cutting distance, a pronounced rounding of the groove ground occurs.

Distance (mm)

T520 Downmilled

T560 Upmilled

Downmilled

B Upmilled

Downmilled

T44 Upmilled

55

295

535

Fig. 9-16 Groove sections for different milling distances and different material states

Downmilled

Upmilled

238

9 Material States and Surface Conditioning for Mold Inserts

The specimen tempered at 560 8C and the bainitic specimen also show significant form errors; however, they are much smaller than those found for state T520. According to a similar hardness, tool wear and resulting form errors are comparable for these heat treatments. For all heat treatment states, burrs can be observed, but their forms and sizes differ significantly. The burrs on the down-milling side (left in the micrographs) are always larger than those on the up-milling side (right in the micrographs). Generally, a significant reduction of the burr size can be found with decreasing material hardness. The state T520 shows burrs which are so large they cannot be allowed in a microinjection mold, e.g. for reasons of tightness. A reliable deburring without form error increase would be very difficult in this case. For both other states of X38CrMoV5-1, the burrs are significantly smaller, especially on the up-milling side. Even for long milling distances, only short burrs with a narrow base are found. On the up-milling side, it can clearly be seen that burr formation is lowest for the bainitic specimen, although the hardness is similar to T560. Mainly the burr bases are much narrower, which facilitates deburring. For micromolds, these two states show acceptable burr sizes, but only if the mold edges are finished by up-milling. Toolox 44, where significant burrs are only observed for large milled distances, shows the lowest tendency for burr formation. 9.3.5 Influence of the Process on the Material State Close to the Surface

Fig. 9-17 shows the modification of the microstructure close to the surface on the down-milling side for all three heat treatment states after a distance of 535 mm. In state T520, the band-shaped microstructure at the surface indicates heavy plastic deformation during the milling process. Thick white layers, which can only partially be etched, are present in the deformed material. In state T560, a preferred orientation of the microstructure can only be seen in the area around the burr, which experienced the strongest plastic deformation. A thin white layer with a thickness of about 2 lm is observed on the groove flanks and around the inner edge. In the austempered state, the form of the bainite needles in the groove ground indicates strong plastic deformation. A thin white layer with a thickness of about 1 lm is formed in the groove ground. On Toolox 44, no microstructural modifications were detected apart from the plastic deformation of the burr material. The plastic deformations observed with the optical microscope lead to a hardness increase as shown in Table 9-8. The hardness was measured close to the surface and compared with the hardness in a nondeformed area of the same former austenite grain. As could be predicted from the microstructure analysis, state T520 shows the largest hardness increase of about 300 HV, the hardness increase of the other states investigated being much lower. This indicates better cutting of these materials.

9.4 Surface Conditioning of Micromilled Structures in a Tool Steel T520

T560

B

T44

Fig. 9-17 Microstructure close to the surface after a milling distance of 535 mm (down-milling side) for different material states

Table 9-8 Hardness increase through micromilling

State

Surface (HV)

Bulk material (HV)

T520 T560 B T44

1060 734 718 529

759 630 633 447

The bainitic state and Toolox 44 show the best machinability of the states investigated. Toolox 44 shows a lower tendency for burr formation, which is due to its lower hardness.

9.4 Surface Conditioning of Micromilled Structures in a Tool Steel 9.4.1 Process Overview

Milling is a very common manufacturing process for injection molds, as it allows fast and economic fabrication of these molds. As shown in the previous sections, surfaces made by milling usually show striations in their topography which result from the cutting process. When milling hard material states, burrs appear on the mold edges. In injection molding, these surface states lead to bad removability of parts with a low flank inclination from the mold. Burrs furthermore cause deterioration of the tightness between mold parts, leading to flashes on the workpiece in the joint level. For these reasons, a direct use of milled injection molds is often not possible and smooth finishing is necessary. For this purpose, usually manual grinding and polishing tools are used. As a manual smooth finishing is not possible with micropieces, alternative processes must be found for micromold manufacture. In precision engineering,

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9 Material States and Surface Conditioning for Mold Inserts

nonmanual fine machining processes are well known. For use in micromanufacturing, these processes must be modified. Processes for smooth finishing can be divided into three classes: · peening processes; · laser-based processes; · electrochemical processes. According to the blasting agent used, peening processes can be divided into abrasive processes, hardening processes and polishing processes. For abrasive processes, fine abrasive powders such as corundum or silicon carbide are used. They are available in grain sizes down to 2 lm [16]. Abrasive peening allows the removal of surface layers or burrs. For hardening peening, mainly glass or zirconia beads with sizes down to 10 lm are used. Hardening processes generate compressive residual stresses close to the surface by inhomogeneous plastic deformation. Surface roughness can usually be reduced by leveling of profile peaks. For polishing processes, vegetable blasting agents such as ground walnut shells or apricot pits can be used [17]. For mold and die production, a combination of these three types of peening processes is often used. In a first step, burrs and striations are removed by abrasive peening; in a second step, the surface is leveled out by nonabrasive micropeening; and in a third step, the surface is polished with vegetable blasting agents. By such a three-step peening process, roughnesses Rz smaller than 0.7 lm can be obtained in hardened steels (9.4.2). For conventional mold production, abrasive processes with a combined airwater jet are used [18]. As the feed rate, blasting pressure and grain size of the abrasive agent are independent, it is possible to adapt the abrasion rate within wide ranges to the material state, the workpiece geometry and the layer thickness to be removed [19]. By regulating the working parameters, a constant surface roughness can be obtained. The compressive residual stresses generated by this type of process have positive effects on the lifetime of molds and dies. Furthermore, process automation is possible. However, the process applicability for microstructures has not yet been established. In dental prosthesis manufacture, peening processes are used for cleaning of metallic and ceramic dental prostheses [20, 21]. Bead sizes down to 25 lm are used, which could be too coarse for applications in microtechnology. Hence peening machines for dental applications have to be modified for the use of finer blasting agents. A laser-based smooth finishing process has been developed at the Fraunhofer Institute for Laser Technology (ILT) in Aachen. In this process, the metal is superficially melted in a depth of about on 1 lm by an Nd : YAG laser, reducing roughness and surface texturations resulting from preceding manufacturing processes. Burrs and pronounced roughness peaks can be ablated. The regulation of the laser parameters ensures that surface deteriorations such as microcracks are avoided [22]. Other manufacturing steps can be integrated in this process, e.g. surface structuring, dimensional corrections or hardening of weak edges. This laser process is now offered commercially [23].

9.4 Surface Conditioning of Micromilled Structures in a Tool Steel

Electrochemical treatments are used for deburring of complex structures and hardly accessible geometries, e.g. gear parts in the automotive industry [24]. At the same time, surface textures can be removed and the roughness can be reduced. Galvanic deburring is based on the fact that the current density is the highest at the edges. Hence burrs with a slim base and a high surface-to-volume ratio are the easiest to be removed [25]. With a convenient geometry of the backing electrode, the material removal can be limited to a small area, thus limiting allowance overstepping through general metal removal [26, 27]. Electrochemical deburring of micromilled grooves made of stainless steels has been successfully performed [28]. The Isotropic Finishing Process developed by REM Chemicals is a combined chemical and mechanical process. The workpieces are treated in a vibrating tumbler with an oxalic acid-based pickle and alumina burnishing balls. The pickle transforms the workpiece surface into a soft and chemically tight coating that protects from further chemical metal removal. On the profile peaks of the workpiece’s surface, the layer is removed by the burnishing balls and the blank metal is transformed again by the pickle [29]. Hence roughness peaks are selectively removed, leading to a smooth and isotropic surface [30]. Roughness values Rmax [3] lower than 0.4 lm can be achieved. This process is mainly used for finishing of bearing treads and gearwheels, where it reduces break-in times and wear [31] and provides an increase in the dynamic strength [32]. For the finishing of micromolds, this type of burnishing process seems only conditionally applicable, as the contact between the burnishing ball and the mold would be only partially guaranteed and significant edge rounding would be expected. In the SFB 499, wet and dry micropeening processes were chosen for smooth finishing of micromold inserts for the following reasons. As peening processes do not require form tools that could only be used for one type of mold, precision problems in tool manufacture can be avoided. By variation of the peening parameters and the peening agents, the process can be adapted to a wide range of materials and surface states. The use of NC machines allows precise processing of geometric elements such as a mold edge with a burr or a mold flank. 9.4.2 Micropeening 9.4.2.1 Specimens and Experimental Setup The specimens for micropeening were microcavities as shown in Fig. 9-18. They were obtained by micro-end milling of quenched and tempered X38CrMoV5-1 with a hardness of 52 HRC (T520; cf. Section 9.3.1 for the heat treatment). The prismatic part of the cavity, that is milled with the same milling parameters as used for steel micromolds in SFB 499, is essentially used for deburring experiments. The straight groove which is machined by a single tool pass is mainly used for investigations on the roughness development due to peening processes because its surface state is homogeneous over the whole length.

241

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Fig. 9-18 Geometry of the micromilled cavity used as micropeening specimen

An SEM view of a milled specimen is shown in Fig. 9-19. Burrs with a height of about 10 lm can be seen on the upper edge. The milled surfaces show striations characteristic of the manufacturing process. The peening machine used was of the type IEPCO Micropeen 200 with a nozzle diameter of 800 lm. The blasting shots used consisted of a mixture of alumina, silicon carbide and glass beads with a grain size between 10 and 20 lm for abrasive blasting and glass beads with a diameter between 20 and 30 lm for nonabrasive peening. The peening pressure was 2 bar for abrasive peening and 3 bar for non-abrasive peening.

9.4.2.2 Deburring By abrasive peening, the burrs shown in Fig. 9-19 could be totally removed, as can be seen in the SEM picture of the specimen after blasting in Fig. 9-20. For

Fig. 9-19 SEM view of burrs on a micromilled cavity

9.4 Surface Conditioning of Micromilled Structures in a Tool Steel

Fig. 9-20 Deburred edge of a milled specimen after abrasive blasting

that purpose, the nozzle was inclined by 758 to the vertical; the peening direction was perpendicular to the cavity edge. As can be seen in the metallographic section in Fig. 9-21, the edge radius after blasting is smaller than 5 lm after total removal of the burrs.

9.4.2.3 Surface states Fig. 9-22 shows a photorealistic shading of the surface topography of the ground of a microgroove (cf. Fig. 9-14) after milling and after two-step micropeening. The surface textures due to the milling process have been completely removed and the peened surface shows an isotropic topography without sharp striations, which facilitates demolding of injection molded workpieces. With the micropeening process applied, the roughness Rz increases from 0.33 to 0.7 lm, which still is very low and is comparable to the roughnesses usually obtained by end milling. Although a roughness increase occurs owing to the peening process, surface textures that obstruct demolding are removed and burrs can be totally removed without inducing inadmissible form errors. Increased surface qualities can be expected from optimized peening parameters, modified blasting agents or a further finishing process step, e.g. ultrasonic wet peening, described below.

243

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9 Material States and Surface Conditioning for Mold Inserts

Fig. 9-21 Metallographic cross section of the edge shown in Fig. 9-20

Fig. 9-22 Surface topography of a milled (left) and a peened (right) surface (scale: 190 ´ 320 lm)

9.4.3 Ultrasonic Wet Peening

Ultrasonic wet peening is a new process for smooth finishing of microstructures. It was developed at the Institute for Materials Science and Engineering I at the Universität Karlsruhe (TH). In this peening process, the workpiece is immersed in an aqueous suspension of an abrasive agent in which cavitation is generated by an ultrasonic sonotrode. The process is based on the mechanism shown in Fig. 9-23. The abrasive particles concentrate in the cavitation field and act as nuclei for the formation of cavitation bubbles [33]. These bubbles become rapidly unstable

9.4 Surface Conditioning of Micromilled Structures in a Tool Steel

Fig. 9-23 Particle acceleration by cavitation bubble collapse

and collapse by forming a microjet [34], which accelerates the particles. The working condition for the process is that the material removal rate by abrasion must be much higher than the material removal due to direct cavitiation. This implies the use of abrasive particles with very rough surfaces that can act as nuclei for the cavitation.

9.4.3.1 Specimens and Experimental Setup The specimens for ultrasonic peening were cuboids with dimensions of 30 ´ 10 ´ 10 mm. The material states investigated were Armco iron with a polished surface and X38CrMoV5-1 (52 HRC; cf. Section 9.3.1) with a ground surface. Armco iron was used for determining the influence of the process on the material state as it shows a simple single-phase microstructure. X30CrMoV5-1 was used for investigations on surface roughness as it is the material actually used for mold manufacture in SFB 499.

245

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9 Material States and Surface Conditioning for Mold Inserts

Fig. 9-24 Ultrasonic wet peening device

The peening device used is shown in Fig. 9-24. It is based on a Hielscher UIP-500 ultrasonic processor which is entirely computer controlled. The computer is also used for measuring the ultrasonic power delivered and controlling the suspension temperature, which is kept constant at 20 8C by a radiator coil. A titanium sonotrode with a diameter of 34 mm has been used. The peening medium was an aqueous suspension of 49 wt.% of Al2O3 (Alcoa CT-3000 SG, 0.7 lm mean particle size) and 2 wt.% of dispersant (Dolapix CE64). The delivered ultrasonic power was kept constant at 250 W.

9.4.3.2 Experimental Results On polished Armco iron, a peening of 14 h duration leads to a roughness increase by abrasion from Rz 0.18 to 1.1 lm. The microhardness increases from 120 to 160 HV. These two effects of abrasion and hardening show that in this process, abrasive peening is the determining effect that widely dominates other effects such as direct cavitation damage or corrosion. On ground X38CrMoV 5-1, the roughness development was measured during a peening time of 22 h. The results are given in Fig. 9-25. After a peening time of about 16 h, a roughness minimum is reached, where the roughness Rz has been reduced from 1.3 to 0.56 lm. For longer peening times, a strong roughness increase can be observed. This means that for one type of workpiece, the optimal peening time must be experimentally determined to optimize the polishing effect. Fig. 9-26 shows the surface topography after grinding and after 16.5 h of ultrasonic peening. The surface texture due to the grinding process still remains after peening, but the profile peaks have been significantly reduced. With a roughness reduced by half and Rz values of less than 0.6 lm, ultrasonic peening allows one to obtain highest surface qualities. A further reduction in

9.4 Surface Conditioning of Micromilled Structures in a Tool Steel

Fig. 9-25 Roughness development vs. peening time

Fig. 9-26 Topography of a ground (top) and of a peened (bottom) surface

the roughness can be expected if larger particles in the abrasive suspension are used [35]. A combination of wet and dry peening processes is expected to bring further progress in smooth finishing of micromilled structures. In a first step, burrs

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9 Material States and Surface Conditioning for Mold Inserts

and surface textures due to the milling process could be removed by abrasive dry peening without creating excessive form errors. A second peening step with microbeads can reduce the roughness generated by the abrasive peening. Ultrasonic wet peening can finally reduce the roughness to values that can hardly be reached by milling processes.

9.5 Conclusions

The minimum chip thickness and built-up edges were shown to have the greatest influence on the surface roughness in side milling. The minimum chip thickness was found to decrease in a linear manner with the workpiece material’s hardness and strength. It can be estimated from the measured roughness values. Hence, hard material states are to be preferred for reducing the surface roughness. The influence of built-up edges also decreases with the material’s hardness. For softer materials, the cutting speed has the greatest influence on the formation of built-up edges and hence on the surface roughness. The cutting speed should be as high as possible for reaching the best surface states. In very hard surface states, subsurface cracks and large burrs appear, so the usable hardness for good milling results is limited. The burr size was found to decrease with the cutting speed because of thermal softening of the material close to the surface. In end milling, materials with similar mechanical properties show significant differences in their machinability. This is mainly because the machinability is not a material property but a system property of tool, workpiece, process and lubricant described by the roughness, tool wear, form errors and burr formation obtained. Hence it is influenced by many parameters that cannot be simply described by a mechanical property. Furthermore, surfaces obtained by end milling are generated by complex kinematics of the minor cutting edges, so that simple kinematic models such as those used in side milling cannot be applied. From the smooth finishing processes that can be applied to milled microstructures made of steel, peening processes were investigated. It could be shown that micropeening allows reliable deburring of milled microstructures made of steel without increasing form errors. The surface roughness after micropeening remains of the same order of magnitude as the roughness of micromilled surfaces. Thereby surface textures can be removed in a reliable way. Ultrasonic wet peening allows the surface roughness of micromilled structures to be reduced to half of its initial value without generating form errors. Surface textures due to previous manufacturing processes remain but are reduced. Hence combined multi-step dry and wet processes can be expected to optimize micromilled injection molds significantly by combining deburring, removal of surface textures and roughness reduction.

9.6 References

9.6 References 1 V. Hüntrup, Untersuchungen zur Mik-

2 3 4 5

6 7

8 9 10 11

12

13 14 15 16 17

18

19

20

rostrukturierbarkeit von Stählen durch das Fertigungsverfahren Fräsen; PhD Thesis, Karlsruhe, 2000. DIN EN ISO 14577; Berlin: Beuth-Verlag, 2003. DIN EN ISO 4287; Berlin: Beuth-Verlag, 1998. P.-H. Brammertz, Industrieanzeiger 1961, 2, 25–32. U. Walter, Einfluss von Kühlschmierstoffen auf den Zerspanprozess beim Fräsen; PhD Thesis, Karlsruhe, 1998. A. J. Pekelharing, Microtecnic 1960, 10, No. 2, 68–78. A. P. Sokolowsi, Präzision in der Metallbearbeitung; Berlin VEB Verlag Technik, 1955. G. Pahlitzsch, D. Semmler, Z. Wirtsch. Fertig. 1960, 55, 242–247. R. Dornhöfer, VDF-Mitt. 1960, No. 21, 12–26. P. Stockinger, Tech. Rundsch. 1968, No. 40, 25–26. W. Degner, B. Hommel, R. In der Stroth, Fertigungstech. Betrieb 1981, 31, 135–138. N. Ikawa, R. Donaldson, R. Komanduri, W. König, W. McKeon, T. Moriwaki, I. Stowers, Ann. CIRP 1991, 40, 587–594. T. Moriwaki, Int. J. Jpn. Soc. Precision Eng. 1995, 29, No. 4, pp. 287–289. N. P. L’Vov, Machines Tooling, 1968, 40, No. 4, 45–46. S. A. Tobias, Machine Tool Vibration; Glasgow: Blackie, 1965. Mikrokorn; Villach (Austria): Treibacher Schleifmittel, 1997. I. Horowitz, Oberflächenbehandlung mittels Strahlmitteln, Band I; Essen: VulkanVerlag, 1982. K. Weinert, F. Albersmann, M. Finke, G. Guntermann, wt Werkstattstechnik 89, 7/ 8; Heidelberg: Springer-VDI, 1999, pp. 351–355. K. Weinert, M. Finke, M. Schneider, C. Willsch, Präzision im Spiegel; Herbst/ Winter, 1997, pp. 1–8. Gesamtprogramm Dentalgeräte; Winterberg: Harnisch + Rieth, 2000.

21 Renfert GmbH, Hilzingen, www.renfert.-

com 22 Anon., ‘YAG gives steel pieces a polish-

up,’ OLE 2001, December, 11. 23 LFC Laser Finishing Center AG, Herzo-

genaurach, www.lfc-ag.de 24 W. König, Fertigungsverfahren, Vol. 3;

Düsseldorf: VDI, 1990. 25 S. Piesslinger-Schweiger, Metalloberfläche

1984, 38, Nr. 11, 505–511. 26 G. G. Gevorkyan, Y. I. Landau, in: Pro-

27 28

29 30 31

32

33

34

35

ceedings of the 7th International Symposium, Birmingham 1983, Kempton; Bedford: IFS, 1983, pp. 417–431. H. Sauer, VDI-Ber. 1975, 240, 45–50. Th. Schaller, L. Bohn, J. Mayer, D. Scherhaufer, K. Schubert, ‘Mechanically cut microstructure grooves less than 50 lm wide with reduced burr,’ in: H. Reichl, E. Obermeier (eds.), Proc Microsystems Technologies 98, UDE Verlag, Berlin, 1998, 644–666. M. Michaud, US Patent 4 818 333, 1989. M. Salerno, Metal Finish. 2000, 98, No. 7, pp. 8–10. S. Rao, F. Hashimoto, ‘a new rolling contact surface and “no run-in” performance bearings,’ presented at the ASME/STLE Tribology Conference, Maui, Hawaii, 1994. L. Winkelmann et al., Paper 2002-011391, SAE International Off-Highway Congress, Las Vegas, NV, 2002, pp. 39– 51. R. Pohlman, E.-G. Lierke, M. Gregor, Über Ultraschall-Kavitationsvorgänge an Festkörperoberflächen; Cologne: Westdeutscher Verlag, 1964. R. T. Knapp, J. Daily, F. Harnmit, Cavitation; New York: McBrown-Hill, 1970, pp. 343–347. R. Pohlman, H. Kickert, W. Böttcher, Über Beeinflussung und Bearbeitung metallischer Oberflächen in ultraschallerregten Bädern; Technische Hochschule Aachen, No. 1375, Cologne: Westdeutscher Verlag, 1964.

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Advanced Micro and Nanosystems Vol. 3. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31208-0

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10 Microinjection Molding – Principles and Challenges R. Ruprecht, G. Finnah, V. Piotter, Institute for Materials Research III (IMF III), Forschungszentrum Karlsruhe Abstract The main characteristics of microinjection molding are described in comparison with injection molding of macroscopic parts. The view from the macro- to the microrange gives information about the special features of machinery, molding tools and micropatterned mold inserts integrated in the tools. These features change depending on the geometry of the different microcomponents, which will be defined. Some examples are given of microinjection molding of plastic microcomponents and for applications in microoptics and microfluidics. This chapter closes with a state-of-the-art overview of variants on microinjection molding such as in-mold assembly, insert and two-component injection molding. Finally, a brief description of the motivation for the use of micropowder injection molding (MicroPIM) for metal and ceramic microcomponents is given. Detailed descriptions of these MicroPIM variants are given in Chapters 11 and 12. Keywords microinjection molding process; polymers; plastics; metals; ceramics; micropowder injection molding; MicroPIM; microparts

10.1 10.2 10.3 10.4 10.5 10.6 10.7 10.8 10.9 10.9.1 10.9.2 10.9.3

Introduction 254 Types of Microcomponents 255 Characteristic Features of Microinjection Molding 259 Machinery and Peripherals for Microinjection Molding 262 Molding Tools for Micropatterned Mold Inserts 266 Micropatterned Mold Inserts 268 Molding Materials for Microinjection Molding 270 Examples of the Microinjection Molding Process 271 Variants on Injection Molding Technology 273 In-Mold Assembly, Over-Molding, Outsert Molding 273 Insert Injection Molding 274 Injection Molding of Lost Molds for Electroforming 274

Advanced Micro and Nanosystems Vol. 3. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31208-0

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10.9.4 10.9.5 10.9.6 10.10 10.11 10.12

Multicomponent Microinjection Molding 277 Compression Molding/Transfer Molding 278 Micropowder Injection Molding (MicroPIM) 279 Conclusion and Outlook for Microinjection Molding Acknowledgments 280 References 281

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10.1 Introduction

In the injection molding process, a melt is injected into a tool commonly kept at constant temperature [1]. The process starts with the melting of the material in the injection unit. The clamping unit of the injection molding machine closes the tool, which is made of at least two half metal parts. The melt is injected into the closed tool by a piston or a screw of the injection unit acting as a piston. The tool then cools the melt. During this process, the melt is held under pressure by the injection unit to compensate the thermal shrinkage. After solidification of the molding material, the tool is opened by the machine actuator of the clamping unit and the final part is ejected from the tool or picked up by a handling system. After preparing a new portion of melt and closing the tool the cycle starts from the beginning. Fig. 10-1 shows a simple scheme of an injection molding machine with a closed tool. Injection molding is an economical process technology for medium- or largescale series production of near-net-shaped and free-formed components for microsystem technology (MST). The economic success of MST is based at least partly on the valuable contribution of injection molding technology [2–6]. For plastic but also for metal or ceramic microcomponents, modifications of ma-

Fig. 10-1 The scheme shows the main components of an injection molding machine. On the left side the clamping unit opens and closes the molding tool; on the right the injection unit prepares the melt of the molding material and injects it into the closed tool. Inside the tool one or a few mold inserts can be integrated for patterning the cavity

10.2 Types of Microcomponents

chinery, process and tools are necessary to transfer the well-known conventional injection molding process to the special conditions for MST [7–12]. Based on the knowledge of microreaction molding [13–17] and different ceramic process technologies (see Chapter 15), the fundamental research on microinjection molding started in 1986 for LIGA [18–21] (see Chapter 8) microcomponents of plastics and in 1995 for ceramic and metal powders at Forschungszentrum Karlsruhe in cooperation with industry [8, 22–26] (LIGA is the acronym for the process chain of X-ray lithography, electroforming and replication by plastic molding technologies, in German ‘Lithographie, Galvanoformung und Abformung’, for the manufacturing of high aspect ratio micropatterned parts or microparts. Aspect ratio is the ratio of micropatterns’ height to width or detail. The difference between micropatterned parts and microparts, both microcomponents, is defined in Section 10.2). Other research institutes contributed further research activities on these topics [27–33]. Nearly every commercially available thermoplastic – unfilled or filled – [4, 10, 27, 29, 34–36] can be used for the microinjection molding of plastic microcomponents to broaden the plastic material range of MST used for reaction molding [15–17, 37] and for hot embossing [38–43]. In contrast to the injection molding of macroscopic parts, some modifications of the process have to be developed to achieve complete mold filling and damage-free demolding even for patterns down to the submicrometer regime (see Section 10.2). Similar process technology adaptations are necessary for powder injection molding of metal and ceramic microcomponents [23, 24, 44, 45]. For the manufacture of microcomponents made of plastics, metals and ceramics, the Forschungszentrum Karlsruhe runs three modified injection molding machines from 500 to 600 kN clamping force and two special microinjection molding machine systems (for both see Section 10.4), and also two machines for low-pressure injection molding of ceramic microcomponents (see Chapter 12). These machines were developed mainly in cooperation with the suppliers to manufacture micropatterned components for potential applications, e.g. in medical and biological technologies, microfluidics in analytics and chemistry, microoptics for information technologies and analytics, microchemical engineering and microelectronics. The corresponding R&D programs range from fundamental research to feasibility studies and from the production of small series of a few thousand microcomponents to the training of industrial staff. Already several industrial companies use microinjection molding of polymer microcomponents commercially [e.g. 46–55].

10.2 Types of Microcomponents

Machinery and process parameters for microinjection molding are determined by the design and the material of the microcomponents. Microcomponents are divided into three different types: microprecision parts, micropatterned parts and single microparts.

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10 Microinjection Molding – Principles and Challenges Fig. 10-2 From left to right: powder injection molded green compact, de-bindered brown compact and sintered ceramic microcomponent. The example shows a ceramic ferrule with an outer diameter of 2 mm for the connection of optical glass fibers. By grinding, a tolerance of < 1 lm in diameter can be achieved. Source: Forschungszentrum Karlsruhe [57]

The overall dimensions of microprecision parts are often comparable to precision parts in the millimeter up to the decimeter range with one or more special features with tolerances in the micrometer range. An example of a microprecision part is a ceramic connector element for optical glass fibers – a ferrule – about 10.5 mm long with an outer diameter of e.g. 2.5 mm finished with a tolerance of < 1 lm (Fig. 10-2) [56, 57]. In most cases the micropatterned parts have overall sizes ranging from a matchbox up to a postcard with micropatterns on one or several surfaces. CDs or DVDs with nanopatterned surfaces and overall sizes up to the decimeter range are well-known examples. From the viewpoint of injection molding, it is more ambitious to achieve the flatness of a CD than the surface topography. The cycle time for a CD of 15 g of polycarbonate (PC) made by a compression injection molding process in a cooled tool is only *4.1 s, because of the small depth of the micropattern [58]. Advanced micropatterned parts have geometric microfeatures with lengths generally larger than their width. The ratio of length (or height) to width in a micropattern is defined as the aspect ratio. Micropatterned parts for fluidic applications often have aspect ratios near unity (Fig. 10-3) [59], and for optical (Fig. 10-4 [60], Fig. 10-5) or mechanical (Fig. 10-6, see also Fig. 10-23) [61, 62] applications from > 1 up to 5, and in special cases up to 10 or even 30. The microcomponents molded so far range from less than one millimeter to centimeters and possess micropatterns on at least one surface with maximum heights of currently up to 2.6 mm, minimum wall thicknesses of 20 lm and smallest geometric details of 0.2 lm. The ratio of flow length to wall thickness

10.2 Types of Microcomponents Fig. 10-3 Stack of plastic microcomponents for the center part of an electrochemical analysis system (overall dimensions about 4 ´ 4 ´ 4 cm3). On top two micropumps with housings made of injection molded polysulfone (PSU) are integrated in this system [59]. The plastic housing of about 2 mm thickness has fluidic grooves at least 250 lm wide and deep. Source: Forschungszentrum Karlsruhe

Fig. 10-4 Ribcon® connector of plastics for 16 optical multimode glass fibers. Source: Forschungszentrum Karlsruhe, IMT and Spinner GmbH, Munich, Germany [54, 60]

required in the design of the micropatterns was as high as 15 for semicrystalline polymers such as polyoxymethylene (POM) and polyamide (PA12) filled with conductive carbon and reached 20 for unfilled amorphous polymers such as pure poly(methyl methacrylate) (PMMA) and > 200 for polyethylene (PE) applied as a coating of inserts made of stainless steel in the molding tool [61]. One example of a micropatterned part is a microspectrometer with an overall size of about 2 cm2 and a microgrid. For the UV and visible spectra the grid is of 120 lm height, with teeth of 2 lm length and 0.2 lm step sizes [50, 63–65] (Fig. 10-5). Microparts are components with an overall size much smaller than 5 mm and a shot weight of usually smaller than 0.1 g of PE. They are injected into single cavities isolated from each other or connected only by the common feeder and runner system. An example of a micropart is a gearwheel made of PMMA with an overall diameter of < 3 mm (Fig. 10-6).

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Fig. 10-5 An infrared microspectrometer system produced by Boehringer Ingelheim microParts, Dortmund, Germany [50, 65]. The optical part of this system (white surface on the left about 20 ´ 10 ´ 2 mm3) is injection molded from PMMA with a grid of about 500 lm height and teeth of 1 lm ´ 3 lm step size rectangular to the substrate. Source: Forschungszentrum Karlsruhe, IMT

Fig. 10-6 Microgearwheel was manufactured by injection molding of polyoxymethylene (PMMA) together with its substrate. On the surface of the 2.8 mm diameter gearwheel the copy of the milling tracks of the mold insert can be seen. Source: Forschungszentrum Karlsruhe, IMF-III

10.3 Characteristic Features of Microinjection Molding

10.3 Characteristic Features of Microinjection Molding

In a conventional injection molding process, a polymer melt is prepared of plastic granulates in an injection unit. The melt is injected into a closed, cooled tool cavity and is frozen to a solid part. The part is ejected after opening the tool, usually by ejection pins integrated in the tool. A molded part consists of the wished part and the feeder and runner system as a copy of the total cavity. The runner channels the plastic melt from the sprue of the injection unit into the tool; the feeder system directs the melt into the cavities. The separation of the parts from the feeder and runner system can be done after demolding or already inside the tool, e.g. by techniques known from opening of a three-plate tool [66]. These technologies are also applied in microinjection molding [67]. In MST, the cavity in the molding tool contains features in the micrometer range. Therefore, the tool technology is adapted to the molding of parts with details of these dimensions. The most important differences between microinjection molding and conventional processing are the evacuation of the cavity and the so-called Variotherm process [8–10] in the injection molding tool. These features are explained in the following. If cavities containing micropatterns that are so small that they cannot be vented in the classical way, i.e. through the parting plane or special bore holes, evacuation of the cavities prior to injection is necessary [9, 15, 68]. Evacuation of the cavity prevents compression-induced defects in the molding material and in the pinhole-like micropatterns of the mold inserts (similar to the Diesel effect). The evacuation technique is similar to the procedure used for liquid silicone rubber (LSR) injection molding. Thus a vacuum pump has to be added to the machine peripherals. A further characteristic feature of microinjection molding is the so-called Variotherm process (Fig. 10-7). This process was developed because of the large surface-to-volume ratio of many microparts or micropatterns leading to fast cooling or even freezing of the injected melts in cooled tools. Despite the fact that polymers have a low thermal conductivity and usually show a ‘self-isolating’ effect, the injected materials would freeze on the molding tool’s wall and the microcavities would not be filled completely. As a consequence of the thin walls and large surfaces of micropatterns compared with their volume, the temperature of the molding materials adapts to that of the mold within milliseconds. Therefore, the micropatterned mold inserts with high aspect ratios are heated to temperatures commonly above the melting-points or even above the no-flow temperatures of the molding materials, which are usually about 30 Kelvin above the melting-points [69]. After mold filling, the tool has to be cooled to temperatures which allow a safe and defect-free demolding of microcomponents. Often the differences between injection and demolding temperatures amount to more than 50 8C. Therefore, the Variotherm process leads to relatively long cycle times (usually in the range of minutes). On the other hand, it is still the only possibility for

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10 Microinjection Molding – Principles and Challenges Fig. 10-7 Tool temperature as a function of time: the curve represents a Variotherm process during an injection molding cycle for a micropatterned part with microfeatures of an aspect ratio of five made of poly(methyl methacrylate) (PMMA): 1 = closing the tool and evacuation, 2 = injection of the melt, dwell pressure to 3 = demolding the micropatterned part. Source: Forschungszentrum Karlsruhe [8–10, 69]

molding micropatterns of high quality and large aspect ratios, especially most LIGA components [19]. Despite longer cycle times, an economical process can be realized by integration of several mold inserts in an array of microcavities within one tool. In addition to the process parameters such as tool temperature and pressure (see Section 10.8) inside the tool and the mold inserts, the molding material plays an important role. Molding melts with lower viscosities are preferred for complete filling, especially for components with wall thicknesses < 20 lm and ratios of flow length to wall thickness clearly > 10 : 1 and up to 200 : 1 [61]. To achieve these low viscosities, the melts in the injection units are heated to the maximum injection temperature allowed by the material supplier and are injected with high pressure, acceleration and speed. The injection velocity, measured by the screw velocity and affecting the velocity of the flow front in the microcavities, is another important molding parameter. Very high velocities in small cavities, however, should be avoided, because the resulting high shear rates lead to polymer degradation or segregation of fillers in the thermoplastics. Therefore, process simulations help to optimize process parameters, a process-compatible design of the microcomponents and correct dimensioning of the feeder and runner system. The available software tools for the simulation of the injection phase reasonably meet the demands for plastic microcomponents (Fig. 10-8) [68–73]. Using the material data measured by one’s self and included in the simulation software, the calculated process parameters, such as pressure and velocity, differ from the corresponding experimental data by not more than 10%, although the software does not take special microfeatures such as the Variotherm process into account. Differences up to 100% were determined, e.g. for filling pressure using high filled melts such as feedstocks for powder injection molding (see Chapters 11 and 12), because the simulation tools do not consider the impact of

10.3 Characteristic Features of Microinjection Molding

Fig. 10-8 Results of the simulation of microinjection molding of a microturbine housing using the software tool MOLDFLOW. The housing has two cylindrical areas, which act as inlet and outlet and which surround a central axis for a gearwheel. The hollow central cylinder of 1 mm length has a wall thickness of 115 lm. The pictures show the melt 1 ms before (left) and at the time of complete mold filling (right). Source: Forschungszentrum Karlsruhe, IMF-III [68]

Fig. 10-9 The scheme gives an overview of a complete microinjection molding cycle of the Variotherm process: closing the mold (1 + 2); evacuation te (2); injection, when temperature is reached (3); switch off the Variotherm heating, switch to dwell pressure (4); cooling time tk and dosing in the injection unit (5); opening the tool (6); eject the part (7); start of Variotherm heating on th (8). Source: Forschungszentrum Karlsruhe

powders or fillers on the flow behavior of the melt (see Chapter 3). Using one’s own experimentally determined rheological, thermal and mechanical data for the feedstock, the differences between calculated and measured values can be reduced to < 20%. So far, no simulation tool is available to simulate the complete Variotherm process up to demolding. Owing to the very small dimensions and the fragile geometries of most microcomponents, all machine movements affecting the microcomponents have to be carried out very precisely. Therefore, very low velocities for the opening of

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the molding tool and for the forward moving of the ejector pins have to be chosen for the demolding step. A scheme of a molding cycle is shown in Fig. 10-9. The process begins with closing the tool (1 + 2). Even before the cycle is started the tool is heated (8, 1–3). While the tool is closed the complete tool cavity is evacuated (2). When the required tool temperature is reached the molten material is injected into the tool. At the beginning of the dwell pressure time the cooling of the tool starts (4). At the end the cooling phase the screw in the injection unit prepares the melt for the next cycle (5). When the ejection temperature of the tool is reached, the tool is opened (6) and microcomponents are usually partly or completely demolded. After ejection the part (7) the tool is heated again (8).

10.4 Machinery and Peripherals for Microinjection Molding

For the manufacture of micropatterned parts or microparts with high aspect ratios, some modifications of the hardware and software of standard injection molding machines are necessary to realize tool evacuation, Variotherm process and slow movement of tool opening and forward movement of ejection pins. Both fully hydraulic (Fig. 10-10) and fully electric (Fig. 10-11) machines have already been applied in MST. Based on standard machines, they were developed for microinjection molding by the machine manufacturer, usually cooperating with research institutes which defined the main specifications and made the first tests. At present this developing process is continuing for special applications and variants of microinjection molding (see Section 10.9). One of the advantages of hydraulic machines is the acceleration of the molding mass in the injection unit to realize small injection times of < 0.5 s. Advan-

Fig. 10-10 Hydraulic injection molding machine with a clamping force of 600 kN and special features for the manufacture of micropatterned parts of plastics, metals or ceramics with high aspect ratios. Source: Forschungszentrum Karlsruhe, IMF-III

10.4 Machinery and Peripherals for Microinjection Molding

Fig. 10-11 Fully electrical injection molding machine with a clamping force of 500 kN and special features for the manufacturing of micropatterned parts of plastics, metals or ceramics with high aspect ratios. Source: Forschungszentrum Karlsruhe, IMF-III

tages of electric machines are their precise and, if necessary, very slow movements of the linear motors actuating the injection unit and the knee lever for the tool movement. Different kinds of microcomponents need different injection molding machines with respect to clamping forces and injection units. All known microcomponents have been manufactured using injection molding machines of 50– 750 kN clamping force and injection units as small as possible. When selecting a suitable machine for a specific microcomponent, geometric information about the single microcomponent, the desired quantity of parts produced and the number of forming nests (= cavities) inside the molding tool has to be taken into consideration. The projected surface of the entire cavity, i.e. the sum of the forming nests, the sprue and runner system and the cavity pressure during an injection period determine the required minimal clamping force and thus the minimum size of the molding machine. Clamping forces of 400–750 kN for micropatterned parts are helpful for achieving high part quality without any flash. In addition, the tools have much larger volume than the cavities because of the heating and cooling elements in the tools. Finally, the overall size of the molding tool determines the distance of the tie bars and, in combination with the clamping force needed, the size of the injection molding machine. In contrast to the clamping units, the injection units are very small. In the injection units, the melt is kept at a high temperature level. Therefore, the time that the melt spends in the injection units should be kept as short as possible to secure a homogeneous melt quality by preventing polymer degradation during relatively long cycle times from 20 s to, especially for micropatterned parts, 500 seconds. The volumes of nests added to the geometry of the feeder and runner system define the shot volume. The shot volume determines the choice of the screw size and thus the injection unit. The material defines the screw ge-

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ometry with its melting, compressing and homogenizing zones and the lead-in for granulates. To achieve complete form filling during a cycle phase of injection and dwell pressure, the injection unit has to feed the molding mass by a pressure of about 20–200 MPa depending on the smallest wall thickness and the overall flow length. Because of their higher thermal conductivity, highly filled molding masses generally need higher flow rates of the molding mass and higher tool temperatures during the injection phase. On the other hand, flow rate and accelerations for powder-filled melts should be selected as low as possible in order to prevent separation of polymers and fillers. In general, the control units of molding machines for MST should integrate all special features of microinjection molding into an automatic process cycle. Currently, several machine manufacturers offer hardware and software tools with embedded subroutines, which take into account the described features. If such programs are not yet available or if even more specialized processes are necessary, the executive program and possibly some hardware features of the control system have to be modified, preferably in cooperation with the machine manufacturer. Microinjection molding usually requires more peripheral equipment than conventional injection molding. Some of these items such as vacuum pumps and temporization units are already available on the market (Fig. 10-12). However, other devices such as the handling systems for microparts have to be adjusted to the specific features of MST [74] (see Chapter 16). For example, if an automatic gripper drops a macroscopic part, it will fall due to gravity. In the case of single microparts with a high surface-to-volume-ratio, the adhesion forces may exceed the gravitational forces so that additional units are necessary

Fig. 10-12 Powerful temporization units at the periphery of a microinjection molding machine for the molding of micropatterned parts with high aspect ratios. Source: Forschungszentrum Karlsruhe

10.4 Machinery and Peripherals for Microinjection Molding Fig. 10-13 Microinjection molding system for the molding of single isolated microparts of plastic, metals or ceramics. In front of the machine a self-made magazine station is shown. Source: Forschungszentrum Karlsruhe, IMF-III

to strip off the microparts. The machine parameters for micropatterned parts mainly determine the overall melt flow in the cavities and only have a small influence on the parameters in the microfeatures of the cavities. This is similar to injection molding of microparts using the mentioned modified standard machines with relatively large injection units supplying mainly the feeder and runner systems with a melt. Therefore, there is only small control of molding parameters for the micropatterns in the cavities. Therefore, instead of modifying standard injection molding machines, several machine suppliers have made efforts in the last few years to develop special small machines for microparts. One example is the Battenfeld Microsystem 50, which offers a special injection and clamping unit and the possibility of integrating a quality control, handling and packaging system under clean room conditions as a mini-environment solution (Fig. 10-13) [74]. Primarily, the integrated injection unit is specifically designed, since a small screw provides the melt, a first plunger feeds a defined volume of melt in front of a second plunger, which injects the melt into the cavity. The plunger diameters are much smaller than the screw diameter, thus allowing significantly lower shot weights of as small as 25 mg of PE. With this special injection unit, the first in–first out-principle is realized for the melt and with the needle-like plungers a small volume of melt can be injected into the cavity with high accuracy in melt acceleration, speed and volume. Other peripheral equipment known from conventional injection molding, such as automatic granulate conditioning, granulate

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feeding, automatic demolding and handling of parts and optical in-line quality control of microcomponents and cavities, are adapted and integrated in this small manufacturing system. Efforts to develop micromachines for microcomponents led to innovative melting units [71] and prototypes of injection molding desk-machines [71, 75]. 10.5 Molding Tools for Micropatterned Mold Inserts

For MST molding tools, the experience of the tool manufacturer in precision engineering is valuable. These tools with a two- or three-plate arrangement usually have a stable frame integrating the guiding pins and a tempering system. Especially for the Variotherm process, the tools are thermally isolated from the machine plates by polymer–ceramic-composites or by grooved metal plates. Inside the tool’s frame, where the guiding tie bars are located, the ejection system is integrated on one side and the sprue bush on the other side of the tool facing the injection unit. The sizes of the counterpart holes of the ejection pins are as small as they can be drilled or ground in steel or hard metal. The geometries of sprue, gate and runner system are chosen according to the overall part size and to the technique for separating the feeder and runner system from the microcomponents after shaping. Inside the tool, the micropatterned mold inserts are integrated, carrying one or several form nests (Fig. 10-14).

Fig. 10-14 View of a microinjection molding machine with an open tool. In the middle the gripper picks up the demolded micropatterned part made of polyamide filled with carbon black (PA 12). The ejection pins integrated in the left side of the tool eject the part. On the right side two mold inserts (26 ´ 66 mm2 surface per insert) are integrated in the injection unit’s side of the tool. In the background on the right the second injection unit of this two-component injection molding machine can be seen. Therefore, both injection units are located in a horizontal L-arrangement. Outside the tool (on both sides of the picture) sensors for tool inspection are installed to control complete demolding of the part. Source: Forschungszentrum Karlsruhe, IMF-III

10.5 Molding Tools for Micropatterned Mold Inserts

To evacuate the tool, temperature-resistant polymer seals have to be placed between every plate, around every ejection pin and around other features surrounding the hollow form. For heating the mold, hot oil or water is pumped through channels of the tool’s plate. To reduce the cycle time, usually electric heaters are integrated in the tools at the back of the mold inserts. In this case the channels in the tool plates contain only liquid at a low temperature level for cooling. The relatively long cycle times of up to 10 min resulting from the Variotherm process can be reduced by improved mold technology [76] and by the use of simulation programs for the optimization of mold heating and cooling and of the injection parameters [23, 34, 69] (Fig. 10-15). In addition, a few approaches to develop innovative tempering methods such as inductive heating of the tool have been developed, such as induction [77]. Therefore, the economic challenge of the Variotherm process will probably be mastered. To realize microcomponents without flash at lapped joints, in addition to high clamping forces and good surface quality of the tool plates, all edges of tool plates and mold inserts, which are part of the cavity, have to be very sharp and without chamfers. To secure exact rectangular movements, additional guiding structures are mounted inside the molding tool. In this case, the mold inserts are movable in the tooling plates to allow lateral deflections (Fig. 10-16). Usually the tool part with the more demanding micropattern is positioned towards the nozzle of the injection unit, to avoid ejection pins, in most cases actuated from the closing unit’s side, being integrated in micropatterned mold inserts. Then demolding of the micropatterns takes place during the first stage of opening the tool. By the ejection pins, by a handling system or a combination of both the part is completely ejected. In contrast to microparts (see Chapter 16), micropatterned parts can be handled with the same or similar grippers as in precision engineering.

Fig. 10-15 Temperature cycle in a molding tool during one cycle of microinjection molding applying the Variotherm process. The curves show the process development for shortening the cycle time over the last decade using different solutions for tool temporization. The cycle time for a special micropatterned component made of PMMA with microfeatures with an aspect ratio of > 5 was reduced from 18 to 4 min. Source: Forschungszentrum Karlsruhe, IMF-III

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Fig. 10-16 The picture shows both sides of a small three-plate injection molding tool. The guiding structures inside the tool secure exact tool movements. The mold insert is moveably integrated to realize deformation-free demolding. Source: Forschungszentrum Karlsruhe, IMF-III

10.6 Micropatterned Mold Inserts

In order to manufacture microcomponents such as micropatterned parts or single microparts, the injection molding tools are loaded with one or a small number of micropatterned mold inserts with a geometry inverse to those of the desired microcomponents. These mold inserts can be prepared using the following MST manufacturing techniques: · Micromilling or microdrilling of brass or aluminum alloys [78, 79] (Fig. 10-17), steel (see Chapter 5 and [80]) or microgrinding and -turning of hard materials [81, 82]. · Microelectrodischarge machining (lEDM) [83, 84] or microelectrochemical milling (lECM) [85, 86] in microdimensions (see Chapter 7). · Wet-chemical [87] or reactive ion etching [88] in silicon followed by electroforming. · Lithographic processes, especially by the LIGA process (see Chapter 8), followed by electroforming in nickel or nickel alloys (NiCo, NiFe), [89, 90] (Fig. 10-18). · Laser patterning (see Chapter 6), for instance of steels (Fig. 10-19) or hard metals [91]. · Combinations of the processes mentioned above [92] (see Chapter 4). Currently, microcutting and other micromechanical technologies similar to the traditional precision engineering are the most widely used and the cheapest technologies for the manufacture of micromold inserts (Fig. 10-17). The cutting

10.6 Micropatterned Mold Inserts

Fig. 10-17 Micropatterned mold insert of brass (30 ´ 66 ´ 15 mm3) manufactured by milling with a diamond shaft tool. The smallest structural width is about 200 lm and the aspect ratio is about 1.5. Source: Forschungszentrum Karlsruhe, IMVT

Fig. 10-18 UV-LIGA mold insert of a nickel alloy (20 ´ 20 ´ 5 mm3). The 12 mainly identical cavities are designed by Rolla MicroSynthetics AG, Grenchen, Switzerland, and manufactured by cooperation between Mimotec, Sion, Switzerland [93] and Forschungszentrum Karlsruhe, Germany, for a microgearwheel with smallest lateral dimensions of about 100 lm and an aspect ratio of more than 3. Source: Forschungszentrum Karlsruhe, IMF-III

tools used for micropatterning and the material of the tool determine the smallest dimensions in the mold inserts that can be realized. Wet silicon etching or plasma-enhanced silicon etching and different lithographic technologies (UV, Fig. 10-18; X-ray, see Chapter 8), all combined with electroforming, are most familiar to the process technologies for microelectronics or microelectromechanical systems. They are more expensive because of the equipment used, but the best microprecision can be achieved for lateral di-

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Fig. 10-19 The mold insert is made of hardened steel (about 54 HRC) by laser ablation. The smallest width is about 200 lm with an aspect ratio of about 1. On the right side another writing strategy was chosen. Source: Forschungszentrum Karlsruhe, IMF-I [94]

mensions and sidewall quality of the micropatterns in tool inserts especially in the LIGA process. The state of the art of laser ablation shows roughness values in the micrometer range on the bottom of microcavities (Fig. 10-19). Therefore, the development is directed at new writing strategies or a mix of laser sources to achieve better surface qualities in particular for hard and wear-resistant materials [94].

10.7 Molding Materials for Microinjection Molding

Injection molding is a shaping process for three-dimensional parts using a mold and making use of the thermoplastic behavior of materials. This condition is characterized by temperatures above the glass-transition temperatures of amorphous or above the crystallite melting temperatures of semicrystalline thermoplastic materials but below their decomposition temperatures. The molding materials, which in the case of filled thermoplastics are called injection molding compounds or composites, are the raw materials for the product. The thermoplastic materials used in MST are either unfilled or filled with powders, fibers and/or additives for coloring or for improved demolding, UV stabilization and other functions. For MST, most thermoplastic materials can be injection molded, preferably those with low viscosity. The material selection mainly depends on the targeted application fields. Using fiber-filled polymers, attention

10.8 Examples of the Microinjection Molding Process

has to be paid to friction within the micropatterned mold inserts. At present, only limited information about material properties for microinjection molding of specific polymers is available [95]. However, there is broad experience in microinjection molding concerning thermoplastics such as PMMA, PMMI, PC, PSU, PE, POM, PA, PEEK, PEI, carbon black and carbon-fiber-filled plastics such as POM-C, PA12-C, POM-CF, PA12-CF, PA6.6-CF, PEEK-CF and several ceramic-filled polymers such as LCP [96, 97]. In order to manufacture metal or ceramic parts, the molding materials – the so-called feedstocks – are filled with considerable amounts of metal or ceramic powders, usually to more than 45 vol.% [25]. Although several feedstocks are commercially available, especially in microtechnology, most manufacturers produce their own compound because very fine powders and binder systems fitted to MST application are necessary (see Chapters 11 and 12). In most cases the feedstocks for MST are more expensive, but usually they are needed in small quantities only. Experience exists with both commercial and experimental feedstocks based on metals such as Fe, 17-4PH, 316L [24], Cu, Au, W [98] and WC– Co [26] (see Chapter 11) and ceramics such as Y2O3–ZrO2, Al2O3, ZrO2–Al2O3, Al2O3–TiN and Si3N4 [25, 99, 100] (see Chapter 12). Molding materials have to be stored in a dry atmosphere. For defect-free injection molding, especially in the microrange, the humidity should not exceed 0.1–0.3%. At the Forschungszentrum Karlsruhe all materials and feedstocks for microinjection molding pass through a drying procedure, e.g. polycarbonate (PC) for about 12 h at 120–130 8C in an evacuatable dry furnace. The materials are granulated with an average particle diameter of about 2.5 mm, in special cases < 2 mm. To avoid cross-contamination, a particular injection unit is recommended for each molding material.

10.8 Examples of the Microinjection Molding Process

To illustrate the differences between microinjection molding of microparts and of micropatterned parts, two examples together with the main process parameters are given. The first example is a tensile strength bar, a micropart injection molded of POM (see Fig. 10-20). This test specimen has an overall length of 4.5 mm and a minimal cross section of 130 ´ 260 lm2. The process was developed on a 50 kN Battenfeld Microsystem 50 machine with an evacuatable tool. The cavity was manufactured by wire electrodischarge machining. The ejector pin inside the cavity has the shape of the cavity itself. The temperature of the plastic melt was 180 8C. The tool was cooled to 40 8C on the side of the injection unit and to 20 ± 1 8C on the side of the closing unit, where the cavity is integrated. The specific injection velocity was 350 m/s. The injection phase was stopped less than 5 s after closing the tool. The time of dwell pressure was < 0.5 s.

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Fig. 10-20 Tensile strength bar injection molded of polyoxymethylene (POM) by isothermal temperature management inside a small tool on a 50 kN injection molding machine. Source: Forschungszentrum Karlsruhe, IMF-III

Fig. 10-21 Housing for a microturbine with an overall diameter of 4 mm injection molded from PMMA. Source: Forschungszentrum Karlsruhe, IMF-III

The second example is a microhousing, a small microstructured part (Fig. 10-21) made of PMMA injection molded on a 500 kN Ferromatic Milacron Electra 50 electrically driven machine. It was manufactured in a heavier experimental tool with several temporization units and sensors integrated. With this tool, the Variotherm process can be realized. Two mold inserts up to a micropatterned surface of 20 ´ 60 mm2 can be integrated. The plastic melt temperature was raised to 235 8C, which is the maximum temperature recommended by the material supplier. Before injecting the melt, the tool was evacuated and heated to 130 8C.

10.9 Variants on Injection Molding Technology

The melt velocity was 200 m/s using a specific injection pressure of 1050 bar. After a 1.1 s injection time, the cooling of the tool starts under a dwell pressure of 800 bar. Within a cycle time of 420 s, the micropatterned part was ejected at a tool temperature of 75 8C. These and similar parts are also used as lost plastic models for the development on a microcasting process as a variant on investment casting (see Chapter 13).

10.9 Variants on Injection Molding Technology

Injection molding is not only a process technology to manufacture microcomponents consisting of a single polymer, but also those made of one metal or ceramic using powder-filled mold materials named feedstocks. In addition to the powder injection molding (see Section 10.9.6 and Chapters 11 and 12) further variants of injection molding have already been tested for utilization in the manufacture of plastic microcomponents [101, 102]. Some variants allow the combination of two or more polymers in one microcomponent [70, 103]. In the future, some of these variants will also be of interest in micropowder injection molding [100]. Some injection molding technologies can be used to reduce assembly effort and simultaneously improve the integration degree of components or systems. Obviously, handling and assembly in MST are much more difficult than procedures in macroscopic dimensions. Therefore, the development of micromanufacturing processes that combine shaping and mounting steps is of great importance. By insert molding, in-mold assembly, over-molding, outsert molding or twocomponent injection molding, first multifunctional components were manufactured as described below. For example, a component can be integrated in a molding tool and covered with a thermoplastic material. The components integrated in the tools can be made of plastic, metal or glass. Experiments with ceramic components are under way. The shape of the injection molded plastic phase is patterned by the cavity of the molding tool. These processes allow a reduction in the extent of assembly of single microcomponents to one system by combining two or more different materials inside a single mold within one process cycle. 10.9.1 In-Mold Assembly, Over-Molding, Outsert Molding

In in-mold assembly, a component of the first material is inserted by a handling system into a molding tool with special cavities for fixing the inserted piece. After closing the tool, a molten mass is injected, filling the cavity. As a result, a part consisting of two different materials is generated. Axles or screw fixings are included in the assembly in the same step as snap fits can be molded. This

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process produces a finished construction with a precise adjustment. The technology of in-mold assembly was examined even for microapplications such as joints and valves and also the over-molding of fibers with a lens structure [104]. Over-molding is the term usually used for the over-layering of a component with further material of different properties yielding parts with improved overall properties or with different colors. An alternative way to apply insert parts would be outsert molding, i.e. to manufacture microparts fixed on much larger substrates. Usually the latter consist of metals, but also polymeric units are possible. They are placed in the molding tool prior to closing. To a certain extent, injection molding on conductive templates (see Section 10.9.3) or the utilization of auxiliary demolding structures [105] can be regarded as variants on outsert molding.

10.9.2 Insert Injection Molding

By insert injection molding, an insert is fixed in the tool, a melt is injected and frozen, partly in contact with the insert. Insert injection molding normally aims at a stiff and durable connection between two shaped materials. Because of tool cooling and the cold insert, the melt freezes faster than without an insert. Especially movable or articulated connections between the two components, both usually made of polymers, provide the integration of mechanical functions in one part. In this case the two polymer components must be chosen carefully concerning their melting-point. They should be of low compatibility, in order to obtain one part with two or more movable components and no fixed conjunction. This application is used for the production of microgears [105]. Different trials of insert injection molding for microcomponents have been carried out at the Institute for Plastics Processing, Aachen, Germany, with the aim of producing, for example, hybrid optical or electrical parts [104]. The technology of insert injection molding has also been investigated at the Forschungszentrum Karlsruhe, where steel inlays were layered with polymer films with a minimum thickness of 15 lm and length-to-wall-thickness ratios of up to 250 [61]. 10.9.3 Injection Molding of Lost Molds for Electroforming

Another special application of the insert technique to the manufacture of metal microparts is under development at the Forschungszentrum Karlsruhe. Injection molding is used for the production of two-component (= 2C) thermoplastic molds for the electroforming of metal microparts with high aspect ratio and surface quality. For the electroforming of defect-free metal microparts, these micropatterned plastic lost molds consist of an electroconductive substrate with insulating plastic micropatterns on top [103] (Fig. 10-22). Electroplating (see Chapter 14) of a large number of metal microparts on a single thermoplastic lost

10.9 Variants on Injection Molding Technology

Fig. 10-22 Insert injection molded two-component microcomponent with insulated polymer micropatterns made of polyoxymethylene (POM) (white) on an electroconductive plastic substrate of polyamide PA 12 (black). The micropattern’s height is about 300 lm, the smallest wall thickness is 100 lm and the smallest gap is 30 lm. Source: Forschungszentrum Karlsruhe, IMF-III

mold seems to be a promising method for a future large-scale series-production of metal microparts with good surface qualities. To manufacture a lost plastic mold, a substrate plate with a perforation is molded using an electrically conductive thermoplastic polymer in a first, separate, process step. This substrate is then inserted into a second molding tool with another micropatterned cavity. During the second injection cycle, an electrically insulating component penetrates through the perforations in the substrate plate, which has to be adjusted perfectly with respect to the micropatterned mold insert. The melt fills the spaces of the micropatterns of the mold insert and the microperforations in the substrate plate. By this insert injection molding process, micropatterns of insulated polymers are replicated on top of the conductive substrate to achieve microcomponents of two plastics of different electrical conductivity. For the substrate, thermoplastics such as PMMA, PA6.6, PA12, POM and PEEK filled with carbon black or carbon fibers were tested. For the molding of the insulating micropattern, unfilled thermoplastics such as PMMA, POM and PA12 were used. Further development is aimed at manufacturing these lost molds by the 2C microinjection molding process (see Section 10.9.4). Insulating walls within the lost plastic mold serve to enhance diffusion-controlled metal growth during electroforming perpendicular to the substrate surface without defects inside the metal deposition. If both substrate and pattern are made of a conductive polymer or of a complete conductive surface, electroplating deposition is limited to aspect ratios < 5, because no agitation of the electrolyte’s bulk can support the metal deposition in the microcavities and the current density become to zero on the bottom of the microcavities. For higher aspect ratios, blowhole-like defects will occur in the deposited metal micropatterns caused by an inhomogeneous metal overgrow of the plastic apertures [107].

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The thermoplastic materials of different electric conductivity need to adhere well to each other. Otherwise gaps result between the substrate plate and the insulating micropatterns. For metal deposition without previous surface etching, the electrical resistance of the conducting substrate surface has to be < 1 kX. The value of electrical surface conductivity can be adjusted by the injection rate. Lower injection rates cause enough conductive carbon black to be present on the substrate’s surface [106, 107]. In this way, good nucleation of the electroplating deposition can be guaranteed. Subsequently the 2C plastic molds are filled with metal, e.g. nickel of a sulfamate electrolyte. A metal deposition rate of e.g. 12 lm/h takes some hours to produce the metal microstructures. If electroplating is continued until all individual polymer structures are completely covered by a sufficiently thick metal layer, the resulting metal block is used as a magazine to handle the electroplated metal pattern. After trimming the metal block can also be used as a copy of the original mold insert. If needed, grinding the overgrown metal surface away separates the single metal microparts. In first experiments, gearwheels with an outside diameter of 550 lm were produced (Fig. 10-23). The maximum aspect ratio obtained today by this process is 14. Further developments are aimed at the reproduction of mold inserts and the large-scale production of microparts with finer structural details. Apart from gearwheels, this new technology can also be employed for the production of metal nozzles, spirals, springs or components for microelectronic applications. In addition, first experiments on the development of ceramic microparts by electrophoresis deposition (see Chapter 15) have been carried out at Freiburg University, IMTEK, with lost molds of the Forschungszentrum Karlsruhe.

Fig. 10-23 Microgearwheel with a diameter of about 550 lm made of nickel by electroforming in a polymer lost mold made of conductive substrate and insulating micropattern by insert injection molding. Source: Forschungszentrum Karlsruhe, IMF-III

10.9 Variants on Injection Molding Technology

10.9.4 Multicomponent Microinjection Molding

In 2C injection molding, two different molding masses are injected into a molding tool at the same time or in the same process cycle. Multicomponent injection molding is not only a process to combine several different colors. Innovative variants on the injection molding technology open up the possibility of a higher degree of integration in micromanufacturing. Multicomponent microinjection molding is usually performed by using two or more materials which have a good affinity, if strong joints have to be realized, or which show repellent behavior against each other in case of movable joints. Better than insert molding, 2C injection molding increases the integration grade for microcomponents and reduces the assembly effort. A main advantage of the 2C injection molding technology is that the assembly of two different parts with different physical properties in complex microsystems can be realized without any handling or adjustment of components of the parts. Multifunctionality can improve the performance, the quality and the costs of microproducts. Investigations towards 2C injection molding of microcomponents regarding mold ability, adhesion or repulsion have been carried out at IMM Mainz [105] and also at the Forschungszentrum Karlsruhe [61]. Filled and unfilled thermoplastics (see Fig. 10-24) or two different feedstocks have been used for two-component microinjection molding, respectively [100]. In contrast to insert molding, the materials in multicomponent molding are combined in

Fig. 10-24 Two-component microinjection molding of tensile test bars made of polyoxymethylene (POM, min.: 480 ´ 480 lm2) of two different colors. The different locations of the interconnection area are realized by changing the injection molding speeds of the different types of POM. Source: Forschungszentrum Karlsruhe, IMF-III

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the fluidic state for good adhesion or precise interconnection area. Therefore, the technology is also interesting for powder injection molding (see Sections 11.5 and 12.3.8) in so far as combinations of materials with similar sintering regimes and sintering shrinkages can be chosen. One goal of 2C injection molding is to manufacture composite parts combining the different properties of both microcomponents. Furthermore, mechanical or optical or electrical properties (see Section 10.9.3) can be varied by a proper material combination. A flexible polymer material connected with a stiff thermoplastic can produced seals or a different hand feel [70]. Another well-known technique, which combines two materials with different properties, is the 3DMID (three-dimensional molded interconnected device) technology [108, 109]. The goal is to integrate a circuitry on to a two- or three-dimensional surface. A technology to manufacture the conductive circuit pattern integrated in a plastic part is the combination of 2C injection molding with electroforming. One of the two thermoplastic components contains a plating catalyst (e.g. Pd) so that, after a chemical treatment, it can be coated with a metal [110]. With regard to the strength of the composite and also to its adhesion to the metal layers of MID parts, several thermoplastics have been examined [111].

10.9.5 Compression Molding/Transfer Molding

Although the basic macroscopic injection molding technology achieves relatively high accuracies of the plastic parts produced, applications often demand even better precision. Typical examples are optical lenses where slight deviations of the geometry will decrease the optical properties significantly. An additional problem has to be solved, namely how to cool a bulky part whose cooling shrinkage can be compensated by backpressure for a short time only. A solution to both challenges is the so-called compression injection molding technology, whose typical characteristic is an additional forming step after injection of the molding melt realized by the movement of part of the molding tool. A movement of the mobile clamping plate can carry it out with the impulse vector parallel to the machine axis or by additional plunger units within the molding tool [112]. A widely used sequence of this variant of injection molding consists of the following steps: the molding tool is, at first, not completely closed so that the injected melt fills the cavity only partly. Subsequently, a pushing movement of the machine closes the tool and all sections of the cavity are filled. In this way, the moving force of the compression step is delivered by the machine and provides an additional shaping force strictly rectangular to the parting plane of the tool. Further advantages of compression molding are reduced internal stresses, less deviations, prevention of in-part cavities and higher flow length to wall thickness ratios. In the case of semicrystalline thermoplastics an even higher crystallinity might be achieved. One of the best-established applications of compression molding in industry is the fabrication of the DVD [5]. Owing to

10.9 Variants on Injection Molding Technology

these considerable advantages, compression injection molding is attracting more and more interest and industrial utilization. Injection transfer molding (ITM) has been developed at the Institute of Plastics Processing (IKV) of the RWTH Aachen. It represents a combination of both injection and transfer molding. Compared with the use of molds with hot runner systems, the new technology provides a more homogeneous heat distribution, easier adjustability to different molds and also lower costs. It has been tested mainly for durable resins, but also trials with thermoplastics and even ceramic feedstocks have been carried out [113]. 10.9.6 Micropowder Injection Molding (MicroPIM)

A detailed description of micropowder injection molding for metal and ceramic microcomponents is given in the Chapters 11 and 12. The motivation for the use of these variants in microinjection molding is given in this section. Microsystem technology products penetrate all kinds of human life, from household appliances to medical/biological technology and from information technology to industrial applications such as microchemical engineering. It can be easily imagined that this wide range of completely different applications requires various solutions for the particular microcomponents, each of them with its own technological requirements. As a consequence, an even wider range of materials and related process technologies has to be made available to realize such a variety of devices. Currently, silicon and a number of selected polymers and metals dominate the range of materials in MST, mainly owing to microelectronics. Consequently the established micromanufacturing processes are also focused on these materials. To overcome this unsatisfactory situation and to enlarge the portfolio of engineering materials available for MST, several alternative processing technologies have been developed in recent years. To extend the number of materials for MST, in addition to a wide range of polymers, metals (e.g. Chapters 8, 13 and 14) such as steel and tungsten (Chapters 11 and 19) as well as ceramics (Chapters 12, 15 and 20) are now in the center of replication process development [8, 24, 26, 45, 98, 114–117]. If they can be made available to MST, many improvements and new products can be expected from the millimeter to the nanometer range. If for special applications components are subject to mechanical strains – the classical example is microgearwheels, e.g. for watches – steels or even ceramics [118] can provide higher strength and stiffness. Compared with polymer materials, the abrasion resistance of hardenable steel or ceramics is much better. With wear-resistant materials, the manufacture of mold inserts for microinjection molding will be possible [119–121]. In microchemical engineering, materials with high temperature stability and sufficient resistance against hazardous media are required. The advantage of using ceramics and metal alloys for such devices is obvious.

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Another field of application is microoptics and IT products. Here, the relatively high thermal expansion of polymers is a severe limitation to their use. Therefore, outdoor devices could be made of ceramics (see Fig. 10-2 and Chapter 12). Last, but not least, some physical properties are unique to certain materials. Hard and soft magnetic, piezoelectric properties, etc., may serve as examples (see Chapter 15). From these arguments, the demand for metal and ceramic materials in MST is evident. Additionally, the economic aspects have to be taken into account, i.e. industrially viable processes for small- but also for medium- and large-scale production must be developed. Among others, an interesting approach to meet all these requirements is the adaptation of powder injection molding processes well known from macroscopic industrial fabrication to micromanufacturing [114, 116, 122, 123]. Micropowder injection molding combines the advantages of free shaping possibilities, the availability of powders for nearly every material [25, 124] and an economical process for mass production.

10.10 Conclusion and Outlook for Microinjection Molding

Microinjection molding offers an economical process technology mainly for medium- and large-scale series production of microcomponents made of plastics, metals and ceramics. Mainly for plastics the equipment for microinjection molding, including machinery, peripherals, tool and insert technology are either commercially available or are about to be launched on the market. For the propagation of micromanufacturing technology, it is necessary to determine the material properties of injection molded samples in microdimensions in more detail and to continue the development of the injection molding process for special applications. Further research should also be focused on variants of microinjection molding processes such as multicomponent or powder injection molding and developments on process modelling in the microrange (see Chapter 3).

10.11 Acknowledgments

We thank the Deutsche Forschungsgemeinschaft (DFG-SPP 1012) and the Helmholtz-Gemeinschaft (Program Microsystem Technology of the Forschungszentrum Karlsruhe) for their financial support, our partners, namely several companies and universities, for fruitful cooperation and our colleagues in several institutes of the Forschungszentrum Karlsruhe for their valuable contributions mentioned in this chapter.

10.12 References

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cro Parts; J. Polymer Eng., Vol. 21, No. 2–3, 2001, pp. 87–98. Ruprecht, R.; Gietzelt, T.; Müller, K.; Piotter, V.; Hausselt, J.: Injection molding of microstructured components from plastics, metals and ceramics; J. Microsystem Technologies, 8, 2002, pp. 351–358. Piotter, V. Finnah, G.; Greiner, A.; Plewa, K.; Ruprecht, R.; Hausselt, J.: Considerations on the Simulation of MicroInjection Molding; Gordon Research Conf. On CAE in Polymer Processing, Ventura, Cal, USA, March 6–11, 2005. Ganz, M: Microsystem 50 – die wirtschaftliche und prozeßsichere Verarbeitung von Mikroformteilen; FAKUMA Fachsymposium 19. 10. 99, (editor) KATZ, Aarau, Ch, 1999. Michaeli, W.; Opfermann, D.; Gärtner, R.: Ultrasonic Plastification Concepts in Micro Injection; Euspen ’03 Internat. Conf. on Precision Engineering, MicroTechnology and Measurement Techniques and Equipment, Aachen, May 19–20, Proc.-Edt: Weck, M.; Kunzmann, H.: FhG-Institut Produktionstechnologie, Vol. 2, 2003, pp. 355–358. Kemmann, O.; Schaumburg, C.; Weber, L.: MicroMoulding Behaviour of Engineering Plastics; Symp. Design, Test and Microfabrication of MEMS and MEOMS, Paris, F, SPIE Vol 3680, 1999, pp. 464–471, No 0277-786X/99. Schaumburg, C.; Ehrfeld, W.; Schinköthe, W.; Walther, T.; Weber, L.: Injection Moulding of Microstructures with Inductive Heating; Conf. MicroSystem Technologies ’98, Postsdam, D, edt. Reichl. H.: VDE-Verlag Berlin, D, 1998 pp. 679–681. Schaller, Th.; Bohn, L.; Scherhaufer, D.; Schubert, K.: Microstructure Grooves less than 50 lm Wide Cut with Ground Hard Metal Micro End Mills; ASPE 13th Annual Meeting, 1998 Proceedings Vol. 18, 1998, pp. 108–111. Grimm, U.; Müller, C.; Menz, W.; Wölfle, M.: Fabrication of surfaces in optical quality on pretentious tool steels by ultra precision machining-

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Proc. Euspen Internat. Conf. 04, Glasgow, GB, 2004, pp. 193–194. Schaller, Th.; Heckele, M.; Ruprecht, R; Schubert, K.: Microfabrication of a mold insert made of hardened steel and first molding results; ASPE 14th Annual Meeting, 31. 10.–05. 11. 99, Monterey/USA 1999. Tönshoff, H. K., Freimuth, T.; Reichstein, M.: Ultra Precision Grinding of Porous an Non-porous Super Hard and Brittle Materials for Components of Mciro Systems; in WGP Annals, Production Engineering – Research and Development, Vol. 8, 2001. Klocke, F.; Bertalan, C.: High-Precision Hard Turning of Cemented Carbides; Proc. EUSPEN Topical Conf. on Precision Eng., Micro Technology, Measuring Techniques and Equipment, Aachen, D, 19.-20. May 2003, in: (edt.) Voerde: Rhiem Druck, D, pp. 329–332. Koch, O.; Ehrfeld, W.; Michel, F.; Gruber, H.-P.: Recent Progress in MicroElectro Discharge Machining; Part I: Technology; Proc. 13th International Symposium for Electromachining, ISEM XIII, 2001, Bilbao, Espania, Vol. 2, 2001, pp. 737–745. Förster, R.; Schoth, A.; Menz, W.: Micro-Wire-EDM for production of microsystems in steel and ceramics; Proc. of HARMST 2003, Monterey, California, USA, June 15–17, 2003. Gmelin, H.; Allgeier, U.; Staemmler, L.; Rentz, R.: Endbearbeitung von Formeinsätzen für den Mikrospritzguss mittels elektrochemischen Fräsens mit ultrakurzen Spannungspulsen (ECF); Ergebnisse und Leistungen 2003, pp. 16–17, http://www.uni-stuttgart.de/izfm. Förster, R.; Schoth, A.; Menz, W.: Micromachining of steel using ECM with vibrating tool-electrode; Proc. Euspen Internat. Conf. 04, Glasgow, GB, 2004. Köhler, U.; Guber, A. E.; Bier, W.; Heckele, M.: Fabrication of microlenses by plasmaless isotropic etching combined with plastic molding; J. Sensors and Actuators A, 53, 1996, pp. 361– 363. http://www.protron-mikrotechnik.de, 2004

Bacher, W.; Bade, K.; Leyendecker, K.; Menz, W.; Stark, W.; Thomes, A.: Electrodeposition of microstructures: an important process in Microsystem Technology. N. Masuko, T. Osaka and Y. Ito (Hrsg.), Electrochemical Technology: Innovation and New Developments; Kodansha, Tokio, 1996, 159– 189. 90 Thies, A.; Schanz, G.; Walch, E.; Konys, J.: Electrochimica acta 42, 1997, pp. 3033–3040. 91 Pfleging, W.; Piotter, V.; Hanemann, T.: Laser patterning of hard metal alloys for application in microinjection molding; Proc. European Conference on Laser Treatment on Materials, Hannover 1998. 92 Mohr, J.; Göttert, J.; Müller, A.: Microoptical devices based on free space optics with LIGA microoptical benches: examples and perspectives; Preprint SPIE 2783–2706, 1996. 93 http://www.mimotec.ch, 2004. 94 Pfleging, W.; Bernauer, W.; Hanemann, T.; Torge, M.: Rapid Fabrication of Microcomponents – UV-laser assisted prototyping, laser micro-machining of mold inserts and replication via photomolding; J Microsystem Technologies 9, 2002, pp. 67–74. 95 Haberstroh, E.; Brandt, M.: Determination of Mechanical Properties of Thermoplastics Suitable for Micro Systems; J. Macromol. Mater. Eng. 287, 2002 pp. 881–888. 96 Ruprecht, R.; Hanemann, T.; Piotter, V.; Haußelt, J.: Polymer materials for microsystem technology; J. Microsystem Technologies, Vol. 5, Nr. 1, 1998, pp. 44–48. 97 Ruprecht, R.; Gietzelt, T.; Guttmann, M.; Hanemann, T.; Heckele, M.; Müller, K.; Piotter, V.: Abformverfahren für mikrostrukturierte Bauteile aus Kunststoff und Metall; FZKA-report 6423, Forschungszentrum Karlsruhe, D, 2000. 98 Zeep, B.; Piotter, V.; Ruprecht, R.; Hausselt, J.: Metal injection moulding of microstructured tungsten components for heat transfer promoters in a helium-cooled divertor; Junior EuroMat 89

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2004, Lausanne, CH, September 6–9, 2004. Merz, L.; Rath, S.; Piotter, V.; Ruprecht, R.; Hausselt, J.: Micro powder injection molding with advanced materials; Internat. Conf. on the Powder Injection Molding of Matals, Ceramics and Carbides (PIM 2004), Orlando, Flo., USA, March 21–24, 2004. Örlygsson, G.; Finnah, G.; Piotter, V.; Kaufmann, U.; Ruprecht, R.; Hausselt, J.: Powder injection molding and characterization of two-component, electrically conductive ceramic parts; James, W. B. [Hrsg.] Advances in Powder Metallurgy and Particulate Materials; Proc. Int. Conf., Chicago, Ill., June 13–17, 2004 CD-ROM Part 4 S. 113–24 Princeton, N. J.: Metal Powder Industries Federation, ISBN 0-9762057-0-X. Ruprecht, R.; Benzler, T.; Hanemann, T.; Müller, K.; Konys, J.; Piotter, V.; Schanz, G.; Schmidt, L.; Thies, A.; Wöllmer, H.; Haußelt, J. H.: Various replication techniques for the manufacturing of metal microstructures; J. Microsystem Technologies, Vol. 4, 1997, 1, 28–31. Michaeli, W.; Ziegmann, C.: Micro Assembly Injection Moulding for the Generation of Hybrid Microstructures; J. Microsystem Technologies 9, 2003, pp. 427–430. Finnah, G.; Naumann, K.; Holstein, N.; Piotter, V.; Ruprecht, R.; Hausselt, J.: Herstellung von metallischen Mikrokomponenten durch Einlegespritzgießen und anschließende Galvanoformung; J. Galvanotechnik, 95, 2004, pp. 2776–2780. Ziegmann, Chr.: ‘Polymer processing and microassembly’, Ph.D. Thesis, RWTH Aachen, D, 2001. Ehrfeld, W.; Begemann, M.; Berg, U.; Lohf, A.; Michel. F.; Nienhaus, M.: Highly parallel mass fabrication and assembly of microdevices; J. Microsystem Technologies 7, 2001, pp. 145–150. Piotter, V.; Ruprecht, R.; Konys, J.; Hausselt, J.: Injection molding of microstructures using filled and unfilled thermoplastics; Conf. MicroSystem

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Technologies, Potsdam, 17.–19.September 1996. Piotter, V., Holstein, N., Oskotski, E., Schanz, G., Haußelt, J.: Metal micro arts made by electroforming on twocomponent lost polymer molds; Proc. 3rd Euspen Int. Conf., Eindhoven 2003, pp. 367–370. Arnold, M.; Eberhardt, W.; Giousouf, M.; Kück, H.; Munz, G.; Münch, M.; Oprea, M.; Warkentin, D.: Fabrication of an Electrostatic Miniature Valve from Metallized Micro injectionmoulded Polymers; Proc. MicroTec 2003. Ansorge, F.; Rebholz, Ch.; Badstübner, K.; Wolter, J.: Mechatronische Systeme für die Automobilindustrie; J. Kunststoffe Automotive 1, 2004, pp. 39–41. Marshall, D.: ‘MID-life crisis?’; EU plastics news 10, 1996, pp. 30–34. Eberhardt, W.: ‘Verbundfestigkeit von Thermoplasten bei der Zwei-Komponenten-MID-Technik für miniaturisierte Mikrosystemgehäuse’; AIF-Abschlussbericht, Arbeitsgemeinschaft industrieller Forschungsvereinigungen ‘Otto von Guericke’ e.V. (AiF), D, Report FV-Nr. 12120 N, 2001. Roetting, O.; Roepke, W.; Becker, H.; Gaertner, C.: Polymer microfabrication technologies; J. Microsystem Technologies 8, 2002; Springer-Verlag; pp. 32– 36. Koch, M.; Michaeli W.: Injection Transfer Moulding (ITM) – for the processing of Thermoplastics; Proceedings of Injection Moulding 2002; Vol. 3 pages 7.1–7.9; available at Hexagon Holding ApS, DK-1460 Copenhagen, Denmark. Z. Y. Liu, N. H. Loh, S. B.Tor, K. A.Khor, Y. Murakoshi, R. Maeda: Binder system for micropowder injection molding; Materials Letter 48, 2001, Elsevier Science B.V.; pp. 31–38. Ott, J.; Freiermuth: Precursor Ceramics for Microfabrication; Proc. CIMTEC 2002; Florence, Italy, Vol. 2, 2002, pp. 245–252. Rota, A.; Duong, T. V.; Hartwig, T.: Micro powder metallurgy for the replicative production of metallic micro-

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structures; Microsystem Technologies 8, 2002; Springer-Verlag ; pp. 323–325. Garino, T. J.; Morales, A. M.; Boyce, B. L.: The mechanical properties, dimensional tolerances and microstructural characterization of micromolded ceramic and metal components; J Microsystem Technologies 10, 2004, pp. 506–509. http://www.maxonmotor.com: maxon gear made of ceramic, 2004. Hennige, V. D.; Haußelt, J.; RitzhauptKleissl, H.-J.; Windmann, T.: Shrinkage-free ZrSiO4-ceramics: Characterization and Applications; J. European Ceramic Society 19, 1999, pp. 2901– 2908. Ritzhaupt-Kleissl, H.-J.; Binder, J. R.; Klose, E., Haußelt, J.: Net-shape Ceramic Microcomponents by Reaction Bonding; cfi/Ber. DKG 79, 2002 E9– E12. A. Rota, T. V. Duong, T. Hartwig: Wear resistant tools for reproduction technologies produced by micro powder metallurgy; Microsystem Technologies 7, 2002; Springer-Verlag; pp. 225–228.

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R. Ruprecht, J. Haußelt: Micro powder injection molding; Advanced Engineering Materials, 2, No. 10, 2000, pp. 639– 641. 123 K.-H. Zum Gahr, R. Knitter, R. Ruprecht, K. Seemann, J. Aktaa, W. Brocks, H. P. Buchkremer, R. Gerling, P. Grünberg, M. Peters, W. Reimers: Development of materials and processes for high-performance microtechnology components; in: K. Grassie et al. (Eds.) Functional Materials, Vol 13, Wiley-VCH, Weinheim, D, 2000, pp. 467–472. 124 Hausselt, J.; Ritzhaupt-Kleissl, H. J.; Ruprecht, R.; Both, H. von: Micro and nanopowder-technologies for advanced ceramic microcomponents; (invited) MICRO.tec 2003: Applications – Trends – Visions; Proc.of the 2nd VDE World Microtechnologies Congress, München, October 13–15, 2003 Berlin, D [u.a.]: VDE Verl.GmbH 2003, pp. 45–50, incl. CD-ROM, ISBN 3-8007-2791-9.

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11 Micro Metal Injection Molding V. Piotter, L. Merz, G. Örlygsson, S. Rath, R. Ruprecht, B. Zeep, Institute for Materials Research III (IMF III), Forschungszentrum Karlsruhe, Germany Abstract Metal injection molding (MIM) represents an established manufacturing technology in macroscopic fabrication. For adaption to microtechnology, however, some specialties have to be considered. Feedstocks containing fine fractions of metal powders combining low viscosity with high green strength must be applied. For the shaping process itself, the special features developed for thermoplastic micromolding must also be used for micro powder injection molding (PIM). Although debinding and sintering are in some ways easier in micro than in macro dimensions, grain growth and handling have to be carefully taken into account. The process chain of microMIM – the formation of the feedstocks, tooling, process parameters, debinding, sintering and characterization of the microstructures – will be explained in detail. A further interesting field is the development of two-component microPIM, promising new functional microcomponents and lower assembly efforts.

Keywords metal powder injection molding; MIM; feedstocks; steel; hard metals; cemented carbides

11.1 11.1.1 11.1.2 11.2 11.2.1 11.2.2 11.2.3 11.2.4 11.2.5 11.3

Introduction

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Metal Injection Molding (MIM) in Conventional Fabrication 290 MicroMIM Worldwide 292 MicroMIM Feedstocks

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Binders for MicroMIM 293 Powders for MicroMIM 295 Compounding of MicroMIM Feedstocks 298 Feedstocks Containing Cemented Carbides and Tungsten Rheology of MIM Feedstocks 301 MicroMIM: Machinery, Tooling and Parameters 304

Advanced Micro and Nanosystems Vol. 3. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31208-0

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11.3.1 11.3.2 11.3.3 11.4 11.4.1 11.4.2 11.4.3 11.4.4 11.4.5 11.5 11.6 11.7 11.8

Special Machinery Features in MicroMIM 304 Tooling for MicroMIM 305 Process Parameters for MicroMIM 308 Debinding, Sintering and Thermal Treatment of Microparts Made of Steel and Hard Metal 310 Debinding 310 Sintering 311 Hot Isostatic Pressing 312 Metallographic Investigations of 17-4PH steel 312 Cemented Carbides and Tungsten 316 Two-component MicroMIM 319 Conclusions and Outlook 323 Acknowledgments 323 References 324

11.1 Introduction 11.1.1 Metal Injection Molding (MIM) in Conventional Fabrication

In recent years, manufacturing of metal components by powder injection molding (PIM) has been established as an industrially viable process. The reasons for the wide and rapidly increasing use of this technology are various. Before explaining these benefits in detail, a brief description of the process will be given. MIM requires fine powders with typical particle sizes below about 25 lm. The process is illustrated in Fig. 11-1. At the beginning of the process chain, the polymer binder, which provides the flowability and lubricity of the injection mass (= feedstock), has to be formulated. The metallic constituents and the binder are mixed, often in a two-step procedure, e.g. by preblending in a sigma or z-kneader and homogenizing and pelleting on a shear roll compactor or in a twin-screw extruder. The result is a coarse, granular feedstock ready for injection molding with a volume fraction of powder around 45% to more than 60% depending on the type of metal base powder. The feedstock is processed on conventional injection molding machines with oversized die cavities allowing for 16–20% linear shrinkage during sintering. After solidification in the cooled tool set, the components are ejected and separated from the gates, runners and sprues, which are often recycled. One of the most important advantages of MIM is the high economic efficiency when producing medium and large series. This is due to the shaping step carried out by injection molding, whose high suitability for mass production is well proven by its extensive industrial use in polymer fabrication. Additionally, all following process steps such as debinding and sintering can also be automated. Fig. 11-1 shows a scheme of such a processing line incorporating continuously working debinding and sintering furnaces.

11.1 Introduction

Fig. 11-1 Scheme of a continuous MIM production line with the injection molding machine (1) and the debinding oven (2) in the foreground and the sintering unit consisting of preheating zone (3), sintering furnace (4) and cooling zone (5) in the background. Subsequently, magazining of the sintered parts takes place (6). Courtesy of BASF, Ludwigshafen, Germany

Furthermore, MIM permits the manufacture of complex component geometries with the contours desired, i.e. hardly any finishing work is required. Examples of real industrial applications are given in Fig. 11-2. Compared with the classical ways of metal processing such as mechanical cutting with its long sequence of singular steps such as milling, drilling and boring, MIM achieves significant reductions in time and costs. Regarding casting processes, the shaping

Fig. 11-2 MIM has already entered the field of precision engineering, as demonstrated by these jewelry housings made of stainless steel 316L. Courtesy of ARBURG, Lossburg, Germany

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quality is usually better. Conventional metal casting, for example, often requires a subsequent cost-intensive finishing process. Finally, as practically all known metal and hard metal materials are available in powder form, a wide variety of materials are available for PIM. This wide range not only includes typical powder metallurgical alloys such as 316L and 17-4PH, but also heat-treatable steels such as 100Cr6 and 42CrMo6; even soft magnetic alloys such as NiFe can be applied. Concerning non-iron metals, copper, titanium and nickel super alloys have to be mentioned [1]. These examples are only a selection of the still increasing number of materials important for MIM. For the reasons mentioned above, PIM of metal- (MIM) and ceramic-filled (CIM) feedstocks (see Chapter 12) is a well-established technique in industrial practice today. Worldwide, manufacturers of MIM parts achieve *US$ 1.6 billion of sales and more than 6000 employees are working in this production sector. Concerning quantity, US industry leads with *49% of sales whereas the Asian MIM manufacturers hold 26% and Europe 25%. Fields of use of PIM technology, including ceramic applications, comprise components for industrial machines, automotive engineering, household appliances (‘brown and white’ goods), microelectronics and medical engineering. In the last field, for instance, orthodontic components represent an increasing market for PIM parts. Hence it is not surprising that attempts are made to use PIM process technology for the fabrication of microcomponents also [3]. 11.1.2 MicroMIM Worldwide

In contrast to macroscopic technology, fabrication of components by microinjection molding is characterized by several special features (see Chapter 10). One of them is the manufacture of molding tools. Owing to the very small dimensions, these molding tools can often no longer be microstructured by high-precision engineering methods. If structural dimensions of a few micrometers or even submicrometers are required, lithographic (see Chapter 8) or etching techniques have to be applied. For geometries between the high-precision engineering range and the micrometer scale, microstructuring by laser machining (see Chapter 6) or micromechanical machining (milling, turning, eroding, electrolytic machining; see Chapter 5) represents a good alternative. To a certain extent even MIM itself can be used for manufacturing mold inserts [4]. To achieve high-quality results with such mold inserts, a special process protocol is needed. For instance, the complete mold cavity is evacuated directly before injection. If a certain flow length-to-wall thickness ratio is exceeded, the socalled variotherm process has to be applied (see Chapter 10 and Section 11.3). The combination of MIM with the capabilities of microinjection molding represents a very interesting merger. Consequently, there are several R&D activities in microPIM worldwide. At Sandia National Laboratories, CA, plastic forms injection molded with PMMA were used as sacrificial micromolds for shaping of alumina-epoxy slurries [5].

11.2 MicroMIM Feedstocks

Numerous trials have been carried out at Nanyang Technological University, Singapore, mainly using 316L steel, but alumina or PZT ceramics have also been taken into consideration. Usually, the mold inserts applied were made by etching of silicon [6]. As an interesting example of R&D efforts in Japan, the investigations carried out at Osaka Prefectural College of Technology have to be mentioned [7]. Among others, one main topic is the examination of feedstock characteristics, e.g. how metal powder fraction and particle size affect the properties of the green and sintered bodies. In Germany, research work at the IFAM Fraunhofer Institute in Bremen focuses on the microMIM process [8], with two-component technology also being applied (see Section 11.5). The development of microMIM will be explained in the following sections with a focus on the differences compared with the macroscopic process.

11.2 MicroMIM Feedstocks

Compounding the feedstock is the first step of the injection molding sequence, whose influence on the process chain and on the properties of the final parts should not be underestimated. The demands on a good feedstock are in some ways contradictory. For example, the feedstocks have to contain a high amount of powder to reduce sintering shrinkage and enhance geometric accuracy during sintering. On the other hand, the compounding and shaping processes require the lowest viscosities and the powder distribution in the organic matrix has to be as homogeneous as possible. Therefore, the mixing of feedstocks is a challenging procedure, described comprehensively in the following sections. 11.2.1 Binders for MicroMIM

For PIM, the binder is the medium that provides the required properties necessary to shape solid parts from powder material. Therefore, the main function of the binder is the plastification of a moldable mixture containing the metal powder. When the binder is mixed together with the powder, a viscous mass that can be formed under pressure at higher temperatures is obtained. To fill the mold completely and to replicate all structural details perfectly, the feedstock’s viscosity during form filling has to be significantly lower than for macroscopic MIM. The aim is to obtain a homogeneous feedstock without any powder agglomerates and with low abrasive behavior. Furthermore, the binder/feedstock has to have sufficient strength at the demolding temperature, otherwise, it would not be possible to produce stable parts that can be taken out of the mold, debound and sintered. Therefore, the binder must be suited to be easily removed during the debinding process without any distortion, cracks or other damage to the parts. In practice, depending

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on the components used, binder removal is possible by dissolution in water or other, often organic, solvents if there is at least one soluble component in the binder. Other binders require methods such as thermal treatment or chemical degradation. Chemical degradation can be accomplished fairly fast as the reaction only takes place at the binder surface and, therefore, progresses from the outer surface of the part to the inner volume. Thermal degradation has to be carried out slowly to avoid damage to the parts: if the green part is heated up too fast, degradation products from the inner volume would cause high internal pressure, resulting in deformations and cracks. PIM is usually carried out automatically and demolding often takes place just by opening the tool. Therefore, the forces on the green part are often higher than with manual techniques, so that binders with a high green stability are required. To achieve this, the binders for PIM often contain at least one thermoplastic material as a main component. Polymer materials often used are polyoxymethylene (POM) [9–11], polyethylene (PE), polypropylene (PP), polyamide (PA), poly(vinyl alcohol) (PVA) [12] or modified versions of these components [13]. They can be combined with waxes and additives to form the binder. The waxes are used to reduce viscosity, improve form filling and provide better debinding behavior. They can be melted out at low temperatures to obtain an open porosity, accelerating the subsequent thermal debinding of the thermoplastic component. The additives are used, for example, to optimize powder binder interaction, reduce adhesive forces between feedstock and mold and/or to prevent powder binder separation. The application of these additives depends on the material used, the powder type and size and the geometry of the parts produced. For example, a feedstock for the production of small and detailed parts needs an additive to prevent adhesion of the molded part in the tool, whereas powder binder separation may be critical regarding a feedstock using coarse metal powder with high density. Typical and often used additives to optimize the processability of a feedstock are waxes such as beeswax or stearic acid [14–16]. Using at least two-component binders also has the advantage of step by step debinding. The component with the lower melting-point is melted out at lower temperatures without any influence on the high-melting thermoplastic material. The created open porosity is helpful for the subsequent debinding step and, therefore, minimizes the risk of introducing deformation or cracks in the parts. Several so-called ready to mix binders or even ready to mold feedstocks are available on the market. They are designed for a broad range of applications and fit most requirements of macroscopic industrial fabrication. Examples of commercially available binders and feedstocks are Licomont (Clariant, Gersthofen, Germany) [13], ELUTEC® (Zschimmer & Schwarz, Lahnstein, Germany) [12] and Catamold (BASF, Ludwigshafen, Germany) [17]. Whereas the first one is available as a pure binder system and as complete feedstocks (e.g. Inmafeed based on Licomont), the Catamold system is a ready to mold feedstock available with most basic metal, hard metal and ceramic powders. Regarding the binders Licomont (former) and Siliplast, they are based on polyolefin wax and PVA, respectively. Both can be solvent debound either completely (Siliplast) or partly

11.2 MicroMIM Feedstocks

(Licomont) using, e.g., water or alcohol. Remaining binder components are debound thermally afterwards. Although more a problem for ceramic than for metal microparts, the advantage of this two-step procedure is questionable as the small parts may buoy upwards and are not only drifting away from their assigned place, but are also very difficult to handle afterwards. Unloading the partly debound and now even more fragile microcomponents from the bath and placing them on the sintering tray is, if needed, sometimes problematic. Whereas Licomont and ELUTEC® are based on solvent/thermal debinding strategy, the Catamold system is debound catalytically. The main component POM is decomposed to formaldehyde in a furnace using nitric acid as catalyst. Owing to the naturally higher strength of POM, the green stability of molded Catamold parts is higher compared with feedstocks produced from the binders considered. The relatively high viscosity of the Catamold system has to be compensated by higher tool temperatures. This can be accomplished by, e.g., applying variothermal temperature control (see Section 11.3). For the production of microparts, the range of available commercial binders and feedstocks is limited. They are designed for general use and for the production of consumer goods, normally parts in the centimeter or millimeter range, optimized for short cycle times and easy post processing. As microparts have a much higher surface-to-volume ratio, parameters such as green strength and easy demolding are crucial. For application in microparts, commercial binders (except Catamold) often provide a green stability too low for demolding without damaging the parts. The demand for specially designed binders and feedstocks with fine-grained powders increases with the miniaturization of the parts produced. Therefore, it is necessary to prepare customized binders for the desired microMIM feedstocks by identifying suitable thermoplastic components and waxes and to mix them with the given powder either simultaneously or after a separate step of binder compounding. For some applications, the usage of enforced commercial binders is possible. In this case, the commercial binder is mixed with a thermoplastic component to raise the green stability of the molded parts. The application of such modified binders depends strongly on the possibility of combining an easier demolding step caused by the higher green stability with a deformation-free debinding procedure. Furthermore, the possibly decreasing flowability of the feedstock has to be considered as the viscosity not only depends on the powder particle size but also correlates directly with the viscosity of the pure binder [18]. Typical raw materials for binder development are shown in Fig. 11-3. 11.2.2 Powders for MicroMIM

Metal powders such as stainless steels 316L, 17-4PH, and iron usually have particle sizes from above 1 lm to more than 10 lm. If necessary, a base powder is blended with alloying powders although fully prealloyed powders are also often

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Fig. 11-3 Assortment of thermoplastic (left) and wax components (right and upper row) for binders

used. Compared with ceramic powders with a particle size usually in the submicrometer range, the resulting ratio of particle surface to particle volume is much lower for metal powders, allowing high solid loadings of at least 60 vol.% or even more without causing too high viscosities. In comparison, the solid loadings of feedstocks for high-pressure powder injection molding produced with fine-sized ceramic particles (d50 around 1 lm or even smaller; see Chapter 12) are limited to slightly more than 50 vol.% or even less if nanoscale powders, which often contain hard agglomerates, are used. A material often used in powder metallurgy and also in macroscopic PIM is 174PH hardenable steel. In contrast to most materials investigated for microPIM, commercially available 17-4PH steel powders show fairly large particle sizes. Additionally, the steel powders exhibit a strong tendency for grain growth during sintering. In combination, these facts result in microstructures and surface qualities which are adequate for common engineering parts, but detrimental in the micrometer range. Extensive efforts are being undertaken to obtain fine-grained powders and to elaborate suitable feedstocks and adequate sintering profiles. In Fig. 11-4, the particle size distribution of several 17-4PH stainless-steel powders from different producers (A-types. Carpenter, Richmond, VA, USA; Btypes, Osprey Metals, Neath, UK) is given. Fig. 11-5 shows an SEM picture of the fracture surface of a green part fabricated with a feedstock using metal powder (type A) with 90% of the particles at a size below 22 lm.

Particle size distribution [rel.%]

11.2 MicroMIM Feedstocks

0.01

0.1

Fig. 11-4 Particle size distributions of different 17-4PH stainless-steel powders. Some of these powders are only sporadically available in small amounts and are very expensive, compared with customary grade powders

Fig. 11-5 SEM picture of a fracture surface of a green part fabricated using a feedstock with stainless-steel powder 17-4PH

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11.2.3 Compounding of MicroMIM Feedstocks

The molding material for MIM, the so-called feedstock, is produced by mixing powder and binder. For the production of microparts the use of fine-grained powders is necessary. It is essential that optimum deagglomeration of the powder particles is achieved and that the powder is equally and homogeneously distributed in the organic binder material. Therefore, mixing has to be carried out at temperatures where the viscosity of the binder is already lowered, preferably at temperatures above the melting-points of all binder components. On the other hand, the applied temperatures have to be limited to prevent binder degradation and powder separation resulting from too low a viscosity. Machines applicable for mixing feedstocks especially with fine-grained powders for microapplications are kneaders or extruders. The main differences between these two types of machines are the continuous mode of the extruder compared with the batch mode of the kneader and the higher shear rates generated by the extruder as compared with the kneader. For feedstock production, co-rotating twin-screw extruders or shear roll compactors are used. These are continuously operating machines where powder and binder components are filled in at the feeding zone of the machine. In case of the extruder, the mixing process takes place while the material is transported through the machine and the feedstock is released continuously through the die. To obtain an accurate compound, the use of gravimetric dosing is necessary. The use of an extruder usually leads to several liters of feedstock per hour. During the evaluation process, solid loading tests or the production of small amounts of feedstock are performed with a kneader. Especially for R&D purposes, an instrumented kneader with a small mixing bowl is advantageous. Such machines allow the registration of the material temperature and the applied torque needed to maintain the chosen rotation speed of the rotors. If the volumetric loading of the machine is kept constant, the measured torque directly correlates with the viscosity of the feedstock. Torque diagrams registered with the measuring mixer at Forschungszentrum Karlsruhe during mixing are shown in Figs. 11-6–11-8. Corresponding data for ceramic feedstocks are shown in Chapter 12. The data in Fig. 11-6 show solid loadings from 58 to 72 vol.%, where the critical loading is already exceeded. Furthermore, the torque needed for mixing the feedstocks increases with decreasing particle size owing to the higher surfaceto-volume ratio and the stronger shear forces that have to be introduced to separate the agglomerates of more and smaller particles (see Fig. 11-7). Compared with feedstocks with ceramic powders, the relatively high solid loadings obtained with metal powders directly lead to lower shrinkage because during sintering the thermomechanically driven reduction of the surface energy leads to compaction of the brown body closing those volume sections which were formerly filled by the binder. The lower sintering shrinkage usually results in smaller tolerances, and thus meets better the nominal sizes of the work-

11.2 MicroMIM Feedstocks

Fig. 11-6 Torque data registered with the measurement kneader during feedstock mixing. An increasing powder content leads to higher torque values. If the powder loading exceeds *70 vol.%, no constant torque is achieved, i.e. the mixture cannot be utilized as feedstock for MIM

8.0 7.0

Fig. 11-7 Torque data registered with the measurement kneader during feedstock mixing. If the powder content is kept constant but the particle size is varied, the torque increase with decreasing powder particle size. As an explanation, it can be assumed that there is more binder required to wet the higher specific surface area of the fine powders

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Fig. 11-8 Torque data registered with the measurement kneader during feedstock mixing. This diagram shows the influence of a thermoplastic component in the binder. The torque increases singificantly with a higher content of thermoplastics (here PE)

pieces given by the design. Typical accuracies are in the range of *0.5% (in some cases down to 0.3%) of the nominal dimensions. Nevertheless, from this point of view PIM is still an inferior technology compared with, e.g., milling or drilling or even lithographic processes. On the other hand, the large particle sizes combined with the high densities of the metal powders often cause separation of the powders from the binder during injection molding, especially when a low-viscosity binder with high wax content is used. Therefore, in some cases it is necessary to enforce the binder by adding 10–20% of additional thermoplastic material, accepting disadvantages in rheological properties (higher viscosity) and hampering in a certain way the binder degradation during the subsequent debinding step. Kneading curves of 17-4PH steel feedstocks are shown in Fig. 11-8: the torque increases with increasing PE content. Furthermore, it is not easy to produce finely detailed microparts or microstructured parts with low surface roughness using particles of that size. 11.2.4 Feedstocks Containing Cemented Carbides and Tungsten

Despite the fact that there are numerous PIM manufacturers of hard metal products, only a small number of different WC-Co feedstocks are commercially available. This is due to the great variety of feedstock compositions used by the manufacturers and the small amounts bought. As a result, most of these companies are mixing cemented carbide feedstocks themselves. This allows sophisti-

11.2 MicroMIM Feedstocks

cated feedstock adaptation to achieve the required properties of the hard metal products. The feedstocks used for the experiments at Forschungszentrum Karlsruhe were prepared by premixing tungsten carbide and cobalt powders in different ratios (94 : 6, 91 : 9 and 88 : 12 wt.%) in a tumble mixer for 6–12 h. With these premixed powders feedstocks with solid loadings up to 52 wt.% could be compounded, keeping the mixing torques and viscosities at an appropriate level for microPIM. Another promising materials family to be processed by PIM is tungsten or tungsten alloys. This is due to a few outstanding properties such as strength, hardness, thermal stability or thermal conductivity. Pure tungsten powder is usually produced by precipitation of tungsten oxide, which is reduced under a hydrogen atmosphere to pure metal. Depending on the process parameters such as temperature, atmosphere and even the thickness of the tungsten oxide layers, different particle sizes but also different levels of agglomeration are achieved. In general, the tendency for agglomeration increases with decreasing particle size. The latter is also the reason why tungsten powders are usually subjected to a deagglomeration step prior to compounding. Destroying the hard agglomerates is usually carried out by milling or, in the case of tungsten alloys, during mechanical alloying [19]. For injection molding, mainly W–Ni–Fe, W–Ni–Cu and W–Cu alloys are applied whereas Ag, Re, Ti or Ta are only rarely used. Tungsten powders are not easy to compound and the maximum loading grades mentioned in the literature are *58 vol.% [20]. Alternatively, subsequent compounding of the powder to maximum content can be helpful [21]. 11.2.5 Rheology of MIM Feedstocks

Rheological measurement is an important controlling tool for feedstocks. Especially for feedstock development, the measured data are valuable to improve the compound quality, e.g. binder composition or powder loading. Regarding the development and production of feedstocks, it is especially important for microparts to have feedstocks with low viscosity at higher shear rates, otherwise the mold filling of the small details in the submillimeter range is hindered or even impossible. On the other hand, higher viscosity at low shear rates supports stability during solidification of the green part. The cooling time before demolding might thus be shortened. Because of all this, the selection of binder components is of great importance. To reach the low viscosity values required for proper form filling, one component should be a wax. Unfortunately, with pure wax the green stability for demolding and handling the parts, especially in the case of microparts, is insufficient. To increase green stability, a thermoplastic component is very helpful. Depending on the favored balance between low viscosity and high green stability, the degree of polymerization and the amount of thermoplastic material may vary.

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Normally, the viscosities of pure binders are about 10 Pa · s or lower at molding temperatures and they often show Newtonian or pseudoplastic behavior. Because of these low values, the viscosities of the binders can be measured with several methods such as rotating parallel plate or cone-on-plate tests. For PIM feedstocks, a high-pressure capillary rheometer (HPCR), type Rheograph 2003 (Göttfert Werkstoff-Prüfmaschinen, Buchen/Odenwald, Germany), is used at FZK as it matches the actual range of shear rates and viscosities well. The measurement is carried out by filling the material into a cylindrical chamber which is heated to the required processing temperature. Subsequently the material is pressed through a nozzle with defined diameter fixed on the lower side of the chamber using pressure applied with a tight-fitting piston from the upper side. To obtain progression data it is necessary to measure the viscosity for several shear rates. Data for feedstocks produced with binders of thermoplastic materials and waxes normally show non-Newtonian behavior. With increasing shear rates the resulting viscosity decreases. PIM feedstocks should reveal Bingham or shear-thinning flow characteristics [14, 22]. Bingham flow means that the material does not start to flow until a critical shear stress is applied. Materials show shear thinning behavior if the viscosity decreases with enforced shear rate [23]. The contrary effect, i.e. an increase in viscosity with increase in shear rate, is termed shear thickening or dilatancy. Shear thickening behavior is more likely if coarse powders are used [24]. Furthermore, and in accordance with the results obtained with the measuring mixer during feedstock mixing, the viscosities of similar feedstocks increase with increased solid loading (Fig. 11-9) and/or decreased particle (Fig. 11-10) size at any given shear rate.

shear rate [s–1] Fig. 11-9 Rheological data for a stainless-steel feedstock fabricated with 17-4PH (grain size d90 < 40 lm) measured in the HPCR. The decreasing curves indicate shear thinning behavior for these feedstocks

11.2 MicroMIM Feedstocks

shear rate [s–1] Fig. 11-10 Rheological data for stainless-steel feedstocks with different particle size of 17-4PH measured in the HPCR. The powders were compounded with binders to form feedstocks with a solid loading of 63 vol.%. The resulting viscosities of the feedstocks are in good agreement with the expectations derived from the particle size distribution of the powders

Rheological measurements are difficult for feedstocks with a high solid loading, especially for metal feedstocks. Because of the relatively coarse particles, very high nominal solid loadings can be achieved. When pressed through a nozzle with a too tight diameter, the feedstocks show a slight tendency to separate from the binder at high flow rates. This leads to the extrusion of binder-rich material through the nozzle combined with a concentration of the powder material remaining in the chamber. The result is quasi-dilatant behavior, i.e. increasing viscosity of the materials in the chamber during the measurement, thus falsifying the measured data. Therefore, measurements have to be carried out using a nozzle with a wider diameter to reduce the pressure on the feedstock and the tendency for powder–binder separation. The disadvantage is that high shear rates often cannot be measured, as – owing to the limited chamber volume – the required material for this measurement can hardly be provided.

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11.3 MicroMIM: Machinery, Tooling and Parameters 11.3.1 Special Machinery Features in MicroMIM

Injection molding represents the shaping step of microMIM, resulting in the so-called green compacts. With this technology, it is possible to obtain geometries nearly as complex as in the traditional plastic injection molding [25, 26]. The shaping process in general can be accomplished using standard plastic injection molding equipment, which, however, has to be modified in several respects. The main differences in powder injection molding are relatively high injection pressures (up to 200 MPa), injection parameters adjusted to PIM feedstocks and the use of wear-resistant injection units and screws with an optimized geometry. Therefore, for the successful production of microsized and microstructured parts, the technical equipment has to meet certain demands and some of these are discussed in detail in the following sections. Usually improved wear protection of those sections that come into contact with highly filled feedstocks is recommended. Higher wear resistivity can be achieved by surface treatment of the machinery steels, by hard-plated inline layers or even by manufacturing the whole injection unit in hard metals. The particular choice is an economic one: higher investment costs and longer lifetime for superior wear protection have to be weighed against less expensive standard units which, of course, have to be replaced more often. From the technical point of view, higher wear protection leads to less contamination of the feedstocks, which is especially important for MST if high material purity of the sintered parts is demanded. Owing to the different thermomechanical properties of PIM feedstocks compared with pure polymers, screw geometries have to be adjusted. To avoid plugging effects, the compression ratios are usually reduced to values of *1 : 1.6 or even lower. Additionally, as a result of the higher thermal conductivity of PIM feedstocks, barrel lengths are shorter and should not exceed 18 times the diameter. The injection molding machine should be able to run low shot weights to ensure short feedstock residence times in the hot zone of the barrel. Otherwise, the feedstock suffering from the high injection temperatures often needed for complete mold filling will begin to decompose after a few cycles and the processability worsens. Consequently, the continuous process, fully automated under normal circumstances, has to be run manually or even stopped completely. In terms of the clamping force, the injection molding machine must be powerful enough to withstand the high injection pressures required for complete filling of the microcavities. In combination with a low clamping force, problems like freezing of the feedstock at low injection pressures or overmolding at higher injection pressures will arise.

11.3 MicroMIM: Machinery, Tooling and Parameters

11.3.2 Tooling for MicroMIM

In general, tooling for microMIM is very similar to that of pure or low-filled thermoplastic materials. One important difference, however, is that the cavities have to be oversized to compensate for the sintering shrinkage so that the numerical dimensions can be met as exactly as possible. Compared with thermoplastic microinjection molding, the surface roughness of the molding tool and especially of the microshaped mold inserts should be as low as possible to avoid deviations or fractures of the considerably fragile green bodies. This has to be kept in mind despite the lower thermal expansion of powder feedstocks, which usually results in less problematic demolding forces. For the design of the feedstock-guiding channels and cavities, computer-based simulation tools can provide useful assistance (see Chapter 3). The exact control of the clamping unit’s opening movement and the ejection of the microparts is a further important aspect. To avoid damage to the parts, the opening velocity should be below 1 mm/s until the microparts have left the injection side of the mold. Then the ejector system carefully pushes the parts out of the microstructured cavity where the attached handling system picks them up for deposition on the debinding/sintering tray. The velocity of the ejector pins should be as low as possible to avoid any deformation of the green parts. The position of the ejector pins can also cause deformations in the green part. This problem can be solved by ejector pins shaped according to the contour of the desired part. Fig. 11-11 shows a mold insert with contour ejectors for the production of microtensile test and bending test bars. Examples of microMIM parts manufactured on a laboratory scale using mold inserts produced

Fig. 11-11 Mold insert for tensile (left) and bending (right) test bars with contour ejectors

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Fig. 11-12 Sintered microMIM gear wheel made of 17-4PH steel. The picture shows a part with an outer diameter of *580 lm; the smallest detail is the width of the teeth of *100 lm. This gear wheel was injection molded using a LIGA mold insert

Fig. 11-13 SEM image of a gear wheel fabricated with a micromilled mold insert (see Chapter 5). Outer diameter is *750 lm

11.3 MicroMIM: Machinery, Tooling and Parameters

Fig. 11-14 SEM picture of a sintered microMIM turbine housing part on a substrate plate. Replication was carried out using a LIGA mold insert and 17-4PH steel

Fig. 11-15 The smallest section of the part is the width of the channel *30 lm

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either by micromechanical (see Chapters 4 and 5) or by lithographic methods (see Chapter 8) are given in Figs. 11-12–11-15. Because some of the smallest dimensions of microparts are substantially below 100 lm, the acceptable tolerances are in the range of a few micrometers. Accordingly, there should be no manual treatment of the green microparts until they are debound and sintered to ensure a high and constant quality. 11.3.3 Process Parameters for MicroMIM

In addition to the mechanical aspects, the temperature management of the injection molding tool and the cavity is of fundamental importance. In many cases it is necessary to apply the so-called variotherm method to the injection molding process for microparts. To achieve complete filling of the cavity, the mold is heated to temperatures near the melting-point of the feedstock. After injection of the feedstock, the mold is cooled until the green compact has sufficient strength for demolding. Table 11-1 shows a set of appropriate process parameters. This method is often accompanied by evacuation of the cavity before injection. Otherwise the parts could be damaged by compressed air generating hot spots and thus causing thermal degradation of the feedstock (Diesel effect). Clearly, the procedures described are rather time consuming and might turn out to be a limiting factor for an economical process. For this reason, further development of feedstocks, injection molding machines, molds and mold inserts and also handling systems is essential to keep cycle times as short and the quality of the microparts as high as possible (see also Section 11.1). The molds and mold inserts have to meet different requirements. The most important is surface quality. The surface is replicated by the feedstocks during the injection molding process nearly as exactly as pure polymers can do, which means that nearly every structural detail of the mold will be transferred one-to-

Table 11-1 Typical process parameters for a variotherm MIM cycle

Temperatures of plastification unit: Zone 1 Temperature (8C) 50 Mold temperatures: Injection temperature (8C) Demolding temperature (8C)

130 70

Pressures: Injection pressure (MPa) Dwell pressure (MPa)

80 100

Injection velocity (mm/s)

90

Cycle time (min)

4–5

2 165

3 170

4 175

5 175

Nozzle 175

11.3 MicroMIM: Machinery, Tooling and Parameters

Fig. 11-16 Scheme of a mold insert with microcavities and insertion plate

one to the sintered parts. Furthermore, such details have a strong impact on the demolding capabilities, sometimes leading to disruption of the green parts and clogging of the microcavities with feedstock. This can also happen if the plates of the mold do not match precisely. In consequence, the successful accomplishment of microPIM needs sophisticated mold designs and fabrication technologies, one of which is the LIGA process. Fig. 11-16 shows a mold insert with an insertion plate. This system acts like a simplified three-plate mold tool and allows the direct separation of the desired micropart from the runners and sprues with the opening movement of the tool. As a consequence, subsequent finishing of the sintered parts can be omitted almost completely. Apart from high-pressure processing, with modified binder formulations, alternative molding concepts that apply much lower pressures during shaping were developed and are summarized for ceramic materials with low-pressure injection molding (see also Chapter 12). It has also been shown that injection molding of feedstocks with suitable binders can be processed at intermediate temperatures on die casting machines.

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11.4 Debinding, Sintering and Thermal Treatment of Microparts Made of Steel and Hard Metal 11.4.1 Debinding

After near net shaping by injection molding, the binder has to be removed and the parts must be sintered to obtain dense compacts of the desired material. The removal of the binder is one of the most critical steps of the process, from both the technical and economic points of view [14, 27]. Debinding can be carried out by thermal degradation, solvent extraction or catalytic decomposition of the binder components or by a combination of these methods. Prior to the described debinding steps, a wax or an oily component of the binder could be absorbed by wicking. Because handling of microparts is difficult and the parts have to be removed from the powder bed used for the wicking process and rinsed, the method is disadvantageous in this case. For large parts, the debinding time depends greatly on the wall thickness and limits the economics of the whole process to usually below 5 mm thickness. For microparts with a wall thickness significantly below 1 mm, however, debinding times play only a minor role. Here the interaction of the material with the sintering tray and the conservation of shape are more important than shortening the debinding cycle. Furthermore, one run in the debinding furnace will produce a large number of parts, so the operational capacity is very high even with long debinding times. Examples of thermal debinding and sintering cycles with and without preliminary solvent extraction are shown in Figs. 11-17 and 11-18.

Fig. 11-17 Temperature–time profile for purely thermal debinding of 17-4PH steel feedstock and subsequent sintering in pure hydrogen

11.4 Debinding, Sintering and Thermal Treatment of Microparts Made of Steel and Hard Metal

Fig. 11-18 Temperature–time profile for thermal decomposition of residual binder after partial debinding of the feedstock with a solvent, followed by the sintering procedure (17-4PH steel in pure hydrogen)

The temperature profiles during debinding have to be controlled very accurately to avoid distortion and formation of cracks and bubbles. When, after the initial debinding stage, an open porosity has been achieved, the polymeric component should hold the powder particles together while the temperature is slowly being increased. The part should be completely free of polymer after the first sinter necks have formed. Binder decomposition should be avoided to prevent residues possibly influencing the base materials and the microstructure of the final parts. In the case of carbon development during debinding, iron base materials could be carburized and melt during sintering if the eutectic temperature around 1150 8C is exceeded. To avoid this, the binder can be evaporated under hydrogen or vacuum at lower heating rates. 11.4.2 Sintering

The sintering conditions depend on the type of binder and the metal processed. Batch furnaces running under vacuum or reducing atmospheres are fairly common. More recently, semicontinuous running beam furnaces with separated compartments for debinding and sintering were introduced. Owing to the high specific surface area of the fine powders for metal injection molding and the high sintering temperatures applied, the density of typical components increases to about 95% of the theoretical density or more. During the sintering process, the parts experience a linear shrinkage of 15–22%, depending on the powder loading, the material used and the density reached.

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Fig. 11-19 Pressure and temperature characteristics of an exemplary HIP treatment for steels

11.4.3 Hot Isostatic Pressing

Since the mechanical properties of microparts depend greatly on the size and number of defects, it is desirable to reach a density of more than 98%. For this reason, it can be necessary to apply a thermal or thermo-mechanical treatment such as hot isostatic pressing (HIP) to eliminate residual porosity after sintering [28, 29]. Fig. 11-19 shows the pressure and temperature characteristics of an exemplary HIP treatment for 17-4PH steel. 11.4.4 Metallographic Investigations of 17-4PH steel

Examinations of microPIM of steel powders within the SFB 499 are presently focused on 17-4PH. The methods of feedstock preparation and viscosity measurements have been described in Sections 11.2.3 and 11.2.5. An explanation of the abbreviations ’A‘ and ’B‘ is given in Section 11.2. The discussion of the metallurgical investigations will begin at the sintering step, exerting the most important influence on a material’s microstructure. After debinding, the parts were sintered for 1 h in 100% hydrogen at 1260, 1310 and 1360 8C (Figs. 11-20–11-31). The heating and cooling rate was 3 K/min. At 1260 8C, which is fairly low compared with conventional sintering temperatures for commercial stainless-steel products, the sinter activity was already sufficient to achieve densities of *90%.

11.4 Debinding, Sintering and Thermal Treatment of Microparts Made of Steel and Hard Metal

Fig. 11-20 Microstructure of a microMIM part made of type A 22 lm feedstock and sintered at 1260 8C

Fig. 11-21 Microstructure of a microMIM part made of type B 16 lm feedstock and sintered at 1260 8C

At these densities and as Figs. 11-20–11-23 indicate, there is obviously an open porosity. The initial consolidation of the powders leads to a grain growth by a factor of 2–3. An interesting effect is the formation of a peripheral zone with much larger grains and the explicit change of the microstructure on entering the core material. Only type A 22 lm does not show this effect, probably because of the low sintering activity at this temperature leading to the lowest sintered density of all investigated materials in combination with a large open porosity. A final explanation for this effect, however, cannot be given at present and will be part of further research activities.

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Fig. 11-22 Microstructure of a microMIM part made of type A 9 lm feedstock and sintered at 1260 8C

Fig. 11-23 Microstructure of a microMIM part made of type B 4 lm feedstock and sintered at 1260 8C

On increasing the sintering temperature by 50 8C there is a remarkable change in density and microstructure. Porosity can still be detected in the type A 22 lm powder, but now this porosity is closed. The density ranges between 93.5% for type A powder and 96–97% for the other powders. The formation of a peripheral zone can again be seen in the type B 4 lm powder in Fig. 11-27. Except for these surface layers, all microstructures show a sufficient homogeneity. The average grain size exceeds 35 lm for all materials. Hence one cannot expect to obtain isotropic material properties in microsized parts (see Chapter 19).

11.4 Debinding, Sintering and Thermal Treatment of Microparts Made of Steel and Hard Metal

Fig. 11-24 Microstructure of a microMIM part made of type A 22 lm feedstock and sintered at 1310 8C

Fig. 11-25 Microstructure of a microMIM part made of type B 16 lm feedstock and sintered at 1310 8C

In consequence, further research activities are necessary to optimize the heating and cooling rates or to add grain growth inhibitors to the metal powders. By sintering at 1360 8C, the microstructures of the examined materials again change remarkably. The sintered densities easily reach 97–98%, accompanied by intense grain growth and formation of heterogeneous areas inside the parts. Peripheral zones with a different microstructure cannot be detected. The average grain sizes increase to 50–100 lm, which is too large even for some macrosized

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Fig. 11-26 Microstructure of a microMIM part made of type A 9 lm feedstock and sintered at 1310 8C

Fig. 11-27 Microstructure of a microMIM part made of type B 4 lm feedstock and sintered at 1310 8C

parts. In conclusion, by using fine-grained powders, the sintering temperatures should be kept as low as possible to achieve adequate results. 11.4.5 Cemented Carbides and Tungsten

With the feedstocks prepared as described in Section 11.2.4, different parts such as discs for tribological tests and tensile and bending test specimens were injec-

11.4 Debinding, Sintering and Thermal Treatment of Microparts Made of Steel and Hard Metal

Fig. 11-28 Microstructure of a microMIM part made of type A 22 lm feedstock and sintered at 1360 8C

Fig. 11-29 Microstructure of a microMIM part made of type A 9 lm feedstock and sintered at 1360 8C

tion molded. After debinding, these parts were sintered for 2 h at 1450 8C in an argon atmosphere. Under these conditions, the sintered densities were > 98% of the theoretical densities. The grain sizes in the sintered bodies remained significantly below those of, e.g., commercial products, accompanied by a satisfactory homogeneity of the microstructure. This is shown in Fig. 11-32. In the case of pure tungsten, owing to its high melting-point sintering has to be carried out at high temperatures up to 2800 8C. The sintering procedure can

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Fig. 11-30 Microstructure of a microMIM part made of type B 16 lm feedstock and sintered at 1360 8C

Fig. 11-31 Microstructure of a microMIM part made of type B 4 lm feedstock and sintered at 1360 8C

be accomplished in atmospheres of usually hydrogen or hydrogen mixtures as described by King and Qingfa [30] or as also shown in Fig. 11-33. Even very small amounts of sintering aids have a significant influence on the sintering temperature that has to be applied to achieve an adequate microstructure. For example, with an addition of only 0.4 wt.% of nickel, a density of 99% of the theoretical value can be achieved at a sintering temperature of 1400 8C [31]. In the case of W–Cu alloys, the sintering temperatures further decrease to *1200 8C, whereas W–Ni–Fe alloys are sintered between 1300 and 1600 8C.

11.5 Two-component MicroMIM

Fig. 11-32 SEM microstructure of WC-6Co manufactured with a feedstock developed at Forschungszentrum Karlsruhe especially for MST applications

Fig. 11-33 Microstructure of pure tungsten sintered under an Ar–H2 atmosphere. The feedstock was mixed at Forschungszentrum Karlsruhe

11.5 Two-component MicroMIM

To exploit the full benefits of reduced dimensions in functional microsystems, suitable assembly methods have to be available. These methods are often complicated and time consuming and, therefore, expensive. One way to eliminate some of the assembly steps is the combination of materials with different prop-

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erties during the injection molding process. This should inherently reduce the assembly effort and thus lead to lower costs. For these reasons, two- or multicomponent PIM seems to be a promising method for multifunctional components to be manufactured in medium- or large-scale quantities. The most important technical problem is to keep the materials merged during the sintering process, i.e. the differences in sintering temperatures and sinter shrinkage of both components should not be too high. From this point of view, two-component PIM (2-C microPIM) in microdimensions reveals the advantage that, owing to the much smaller geometries, the internal stresses that occur are lower than in the macroscopic case. In the last 10 years, the combination of different metals using PIM technology has been investigated by several research groups. Co-injection of two materials, both of which flow through the same runner system, was used to produce functionally graded parts with a core and a skin of two different materials [32]. In this way, for example, metallic components with a wear-resistant coating can be produced, saving the additional processing steps otherwise needed for the application of the coating. Metallic components made of magnetic and non-magnetic stainless steels were produced using overmolding in a single tooling [33]: At the start, one part is molded in a first cavity. Next, the mold is rotated, creating a second cavity which is filled using a second injection unit. This injection molding method was also employed in studies investigating the suitability of two metallic materials to be combined via PIM [34, 35]. The compatibility of different metals in a single part was examined using inlays, injection molded in advance, to manufacture two-component test pieces. A paper entitled ’Design guidelines for processing bi-material components via powder-injection molding‘ has been published [36]. Turning again to microPIM, results focusing on two-component microMIM of, for example, ferromagnetic and non-magnetic combinations were reported [37]. Components of 316L and 17-4PH stainless steels, with an interface area of 0.9 ´ 0.9 mm, were manufactured and sintered successfully (Fig. 11-34). Employing a two-step inlay technology on a laboratory injection molding machine, a composite with a sharp interface between the two materials was manufactured. As a potential application for parts of this kind, a magnetic positioning encoder is conceivable. A possible constellation including a magnetic sensoring device is shown in Fig. 11-34. The first steps towards a real two-component microinjection molding process including tool and software development, assembly and functional tests on a two-component machine were carried out at the Forschungszentrum Karlsruhe in 2002–2003. The result was a unique combination of three sub-variants of injection molding: two-component injection molding with powder feedstocks in microdimensions. As these were initial investigations, sample geometries were relatively simple with minimal merging areas of 400 ´ 400 lm and combinations of chemically similar materials such as different kinds of steels or mixed ceramics varying only in their mixing ratio were applied. A view into the tool area of the machine is given in Fig. 11-35.

11.5 Two-component MicroMIM

Fig. 11-34 Test part with ends made of ferromagnetic 17-4PH steel connected by a non-magnetic 316L stainless-steel bar. Micrographic image of one of two boundary regions (right). In combination with a magnetic sensoring device, a potential application will be a magnetic positioning encoder (left). Courtesy of Fraunhofer IFAM, Bremen, Germany

Fig. 11-35 Tool area of the micro two-component powder injection molding machine at FZK. 1, gate of the first injection channel; 2, nozzle of the second injection unit

The 316L and 17-4PH stainless-steel feedstocks employed for the part shown in Fig. 11-34 were also used to manufacture one- and two-component tensile specimens in a two-component injection molding tool designed and operated at the Forschungszentrum Karlsruhe, IMF III. Fig. 11-36 shows an image of the ejector side part of the mold insert for this piece [38]. The cavities of the clamping areas left and right can be seen, and also the testing section and the slot for the sliding split. It is possible to work either with sequential injection of the two feedstocks employing the sliding split to keep one half of the cavity closed

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Fig. 11-36 Ejector side view of a mold insert for a two component tensile strength specimen

during the injection of the first component or with simultaneous injection keeping the cavity undivided through the whole injection molding cycle [39]. In the former case, the second component is injected on to the stationary front of the first component and the position of the interface is, within certain limits, fixed at the position of the sliding split. In the latter case, the position of the interface can be adjusted by varying the injection speed and timing. In that case, tighter joining of the two feedstocks is expected as they are still in a molten state when they combine. The length of the testing section of the tensile strength bars is 12 mm, with a cross-section of 0.48 ´ 0.48 mm. This amounts to aspect ratios (length/width of the narrowest part of the cavity) of 25 and 12.5 for one- and two-component parts, respectively. Fig. 11-37 shows two-component 316L–17-4PH stainless-steel tensile specimens molded in the described tool. The parts were injection molded employing

Fig. 11-37 Two-component tensile specimens made of 316L and 17-4PH stainless steels. Microscopic image of the interface region of a green part, cross-section 0.48 ´ 0.48 mm (left). Green and sintered parts in comparison (middle). Metallographic image of the interface region of a sintered part (right). Courtesy of Fraunhofer IFAM, Bremen, Germany

11.7 Acknowledgments

simultaneous injection of both components. The position of the interface was adjusted by varying the injection speed and timing. Future tasks in the field of two-component microPIM include the extension of the method to hard magnetic materials and hard metals to copper and copper alloys. Additionally, basic research on the rheological behavior and merging flow has to be carried out (see also Chapter 12). This requires comprehensive experiments including feedstock development, tooling, the injection molding process, debinding and sintering. As designs, materials and combinations can be varied over wide ranges, a considerable potential for new developments in microsystems technology is to be expected.

11.6 Conclusions and Outlook

As demonstrated above, microPIM is well on the way to becoming a suitable manufacturing technology for metal microparts. The actual limitations are set by the powder size rather than by the shaping process itself: Improvements in surface quality and replication accuracy demand finer powders in the micrometer range or even below. Process optimization will be carried out in parallel with the development of microspecific simulation tools (see Chapter 3), whereas sub-variants such as two-component microPIM might gain a leading role even for their macroscopic pendants, opening up possibilities to create unconventional material combinations if further basic research is carried out. In general, PIM still is a rather young technology undergoing rapid development with further economically and technically viable variants to be established in the future.

11.7 Acknowledgments

The authors are grateful for financial support provided by the Deutsche Forschungsgemeinschaft (DFG) within the collaborative research project SFB 499. The research work on 2C-micro powder injection molding was conducted with the financial support of the German Ministry of Education and Research (BMBF), grant number 03N1080A-B, in cooperation with the Fraunhofer Institute for Manufacturing and Advanced Materials (IFAM), Bremen. Further thanks go to the large number of cooperating partners, in particular to the Universities of Karlsruhe and Freiburg.

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11.8 References 1 R. Gerling, F.-P. Schimansky, Adv. Eng. 2 3 4 5 6

7

8 9

10 11 12 13 14

15

16 17

18 19

20 21 22 23

Mater. 2001, 6, 387–390. R. Cornwall, www.immnet.com, 2003. V. Piotter, W. Bauer, T. Benzler, A. Emde, J. Microsyst. Technol. 2001, 3, 99–102. A. Rota, T. V. Duong, T. Hartwig, J. Microsyst. Technol. 2002, 7, 225–228. A. M. Morales, R. Pitchumani, A. K. Gutmann, Proc. SPIE 2003, 4979, 430–439. Z. Y. Liu, N. H. Loh, S. B. Tor, K. A. Khor, Y. Murakoshi, R. Maeda, T. Shimizu, J. Process. Technol. 2002, 127, 165–168. K. Nishiyabu, S. Matsuzaki, S. Tanaka, V. Piotter, in: Proc. of PM2TEC 2003, Las Vegas; SPIE: Bellingham, WA, 2003. A. Rota, T. V. Duong, T. Hartwig, J. Microsyst. Technol. 2002, 8, 323–325. J. H. H. TerMaat, J. Ebenhöch, H.-J. Sterzel, Ceram. Mater. Components Engines 1992, 4, 544–551. M. Blömacher, D. Weinand, Keram. Z. 1992, 44, 827–830. M. Blömacher, D. Weinand, Met. Powder Rep. 1992, 5, 43–49. http://www.zschimmer-schwarz.de ? Keramik Elutec, 2004. http://www.clariant.com ? Product ? Licomont, 2004. R. M. German, in: Powder Injection Molding; Princeton, NJ: Metal Powder Industries Federation, 1990. V. M. B. Moloney,. D. Parris, M. J. Edirisinghe, J. Am. Ceram. Soc. 1995, 78, 3225–3232. A. Smolders, J. Sleurs, J. Eur. Ceram. Soc. 1997, 17, 171–175. http://www.basf.de/basf/img/produkte/chemikalien/catamold/Produktblatt_Catamold_e.pdf, 2004. T. Zhang, J. R. G. Evans, J. Eur. Ceram. Soc. 1989, 5, 165–172. P. Suri, S. V. Atre, R. M. German, J. P. de Souza, Mater. Sci. Eng. A 2003, 356, 337– 344. A. Bose, Powder Metall. 2003, 46, 121–126. C. M. Wang, in: Proceedings of PM2TEC Conference 2001; 2001, pp. 8-180–8-192. B. C. Mutsuddy, Proc. Br. Ceram. Soc./ Fabric. Sci. 1983, 3, 117–137. D. Baird, D. Collias, in: Polymer Processing – Principles and Design; New York: Wiley, 1998.

24 J. H.Song, J. R. G. Evans, J. Rheol. 1996,

40, 131–152. 25 V. Piotter, Th. Gietzelt, L. Merz, R. Rup-

26

27 28 29

30

31

32 33

34 35 36

37

38

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recht, J. Hausselt, PM Sci. Technol. Briefs 2002, 4, 18–23. L. Merz, S. Rath, V. Piotter, R. Ruprecht, J. Hausselt, J. Microsyst. Technol. 2004, 10, 202–204. R. Mueller, Landolt-Börnstein New Series VIII/2A1; 2003, Chapter 6. M. Koizumi (ed.), Proceedings of the Third International Conference, Osaka, 1991; 1991. L. Delaey, in: HIP’93, Proceedings of the International Conference on Hot Isostatic Pressing, Antwerp, 21–23 April 1993; 1993. T. L. King, L. Qingfa, in: Processing and Fabrication of Advanced Materials VI; Singapore: Institute of Materials, 1998, pp. 1465–1476. D. M. Won, S.-W. Kim, Y. D. Kim, I.-H. Moon, in: Proceedings of PM2TEC Conference 2001; 2001, pp. 8-173–8-179. J. R. Alcock, Met. Powder Rep. 1999, June, 30–34. L.-K. Tan, R. Baumgartner, R. M. German, in: Advances in Powder Metallurgy and Particulate Materials – 2001, compiled by W. B. Elsen and S. Kassam; Princeton, NJ: Metal Powder Industries Federation, Vol. 4, pp. 191–198. D. F. Heaney, P. Suri, R. M. German, J. Mater. Sci., 2003, 38, 4869–4874. P. Suri, D. F. Heaney, R. M. German, J. Mater. Sci., 2003, 38, 4875–4881. J. L. Johnson, L.-K. Tan, P. Suri, R. M. German, JOM – J. Min. Met. Mater. Sci., 2003, 55, 30–34. A. Rota, in: Advances in Powder Metallurgy and Particulate Materials – 2002, compiled by V. Arnhold, C.-L. Chu, W. F. Jandeska, Jr. and H. I. Sanderow; Princeton, NJ: Metal Powder Industries Federation, 2002, Vol. 10, pp. 49–57. V. Piotter, in: Proc. of Injection Moulding 2003 Conference; Copenhagen: Hexagon Holding, 2003, p. 4. G. Örlygsson, G. Finnah, V. Piotter, U. Kaufmann, R. Ruprecht, J. Hausselt, in: Advances in Powder Metallurgy and Particulate Materials – 2004; Princeton, NJ: Metal Powder Industries Federation, in press.

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12 Microceramic Injection Molding W. Bauer, J. Hausselt, L. Merz, M. Müller, G. Örlygsson, S. Rath, Materials Research III (IMF III), Forschungszentrum Karlsruhe, Germany Abstract Ceramic injection molding is a shaping technology which is well suited for the manufacture of complex-shaped ceramic parts with small dimensions. The great freedom of design and the high degree of automation permit a near net shape and high-volume production and make injection molding attractive for the manufacture of microparts or microstructured devices. The application of injection molding for the manufacture of ceramic microparts requires the development of feedstocks with high flowability and sufficient green strength. In addition to an adapted tooling technology, an enhanced machine controlling is also necessary for the successful shaping of the sensitive features. Owing to high machine and tooling costs, injection molding is normally associated with large-scale production only. However, the broad spectrum of binder systems also provides the opportunity to manufacture small production lots or even a small number of prototypes in an economic way.

Keywords ceramic injection molding; CIM; HPIM; LPIM; feedstock; microsystems technologies; microparts

12.1 12.2 12.3 12.3.1 12.3.2 12.3.3 12.3.4 12.3.5 12.3.6 12.3.7

Introduction 326 Basic Requirements for Ceramic Powders and Binder Systems High-pressure Ceramic Injection Molding (HPIM) 330 Feedstocks for HPIM 330 Binders for MicroHPIM (lHPIM) 331 Compounding the Feedstock 331 Rheology 333 Molding Process 336 Thermal Treatment 338 Example of a Typical lHPIM Production Cycle 340

Advanced Micro and Nanosystems Vol. 3. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31208-0

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12.3.8 12.4 12.4.1 12.4.2 12.4.3 12.4.4 12.4.5 12.4.6 12.4.7 12.5 12.6

Special Development: Two-component MicroCIM 342 Low-pressure Injection Molding (LPIM)

344

Characteristic Features of LPIM 344 Feedstocks for LPIM 345 Compounding the Feedstock 346 Rheology 347 Machines, Tooling 348 Thermal Treatment 350 Examples 353 Conclusions 355 References 355

12.1 Introduction

Injection molding has been used in ceramic manufacture since about 1930 [1, 2]. The driving force was the demand for a large number of spark plugs for automobiles. As the spark plug insulators had a fairly complex shape, which was difficult to manufacture by other methods, a new shaping technique had to be developed to fulfil the demand. Another push came in the 1960s, when injection molding became established in the foundry industry as a routine shaping method for the manufacture of ceramic cores for aircraft turbines. In the 1950s, injection molding was developed for the production of ceramic thread guides. The rapidly growing synthetic fiber and textile industry required a large number of complex-shaped guides which were able to resist the high-speed threads without wear. During the last 20 years, the interest in ceramic injection molding (CIM) has accelerated as the demands of parts design have approached the limits of conventional powder processing routes. Nowadays CIM is a standard shaping method for the mass production of small ceramic parts with complex shapes. Examples, in addition to the already mentioned thread guides, are cutting tools and nozzles for wire bonding machines. However, some frequently presented examples such as turbocharger rotors or turbine wheels [3] are still on a prototype basis and have not yet reached the stage of industrial application as high costs often put up a commercial barrier for the manufacture of large parts [4]. In contrast, CIM plays a commanding role in the manufacture of small or miniaturized devices. These parts are often characterized by a complex geometry which cannot be accomplished by traditional ceramic shaping methods. Shaped by dry pressing or casting, the strength of the unfired (‘green’) compact is low. During the demolding step, i.e. the removal of the part from the mold, there is a high risk that fine details of low-strength materials will be damaged. Mechanical tooling of ceramics is always an expensive issue, even for small work pieces. Owing to the brittle nature of the material, machining can introduce flaws which may lead to failure in use. In contrast, CIM offers a near-net-

12.1 Introduction

shape process for the manufacture of complex-shaped or miniaturized parts with dimensional accuracy. It requires little or no final machining and the high binder content of the injection molded bodies offers sufficient green strength for a damage-free demolding procedure. Injection molding can be fully automated with proper process control, allowing the economic production of large series of parts. The most crucial process step in injection molding, the debinding step, is less problematic for microparts owing to their small wall thickness. However, one has to keep in mind that owing to the thermal process steps, meeting tolerance issues can be problematic. If a microstructure 200 lm in width shows a variation of only 2 lm, the tolerance is 1%. In this respect, the debinding and sintering processes can have a large impact on the functionality of the final parts. As already described in Chapters 10 and 11, the powder injection molding technique comprises three processing steps: feedstock preparation, the molding process itself and the subsequent thermal processing, i.e. debinding and sintering. This is valid for all variants of powder injection molding, micro and macro, and also metal or ceramic injection molding. In this section, the specific features of microceramic injection molding (lCIM) will be described. Two variants of ceramic injection molding exist, the industrially established high-pressure injection molding (HPIM) and the less common low-pressure injection molding (LPIM). In most HPIM processes, thermoplastic polymers are used as the binder component, whereas for LPIM paraffin- or wax-based binders solely are used [5]. In Table 12-1, a comparison between HPIM and LPIM is shown, concerning the characteristic differences in the shaping process. In spite of the different process parameters, both methods generate ceramic microparts with comparable properties (Table 12-2). As HPIM generally requires relatively high investments for machine equipment and tooling, this technique is usually applied for large-scale production running with production numbers > 10 000 pieces. For product development or for smaller production numbers, the LPIM technique is an attractive alternative, because of the markedly lower equipment and tooling costs. Both variants are described below with respect to their capability for the manufacture of three-dimensional ceramic microparts.

Table 12-1 Comparison between HPIM and LPIM: typical shaping properties

Property

HPIM

LPIM

Major binder component Viscosity range (Pa s) Injection pressure (MPa) Processing temperature (8C) Green strength Tool costs Tool wear Units

Thermoplastic polymers 100–1000 > 50 120–200 Medium High High Mass production

Paraffin, waxes 1–20 0.1–1 70–90 Low Low Low Prototypes and small series

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Property

HPIM

LPIM

Green density (vol.%) Maximum bending strength, r0 (MPa) Weibull exponent, m Finest details (lm) Aspect ratio Surface roughness, Ra (nm)

45–55 2100 3.5–10 >2 > 10 possible < 500

50–60 3000 5.6–11.1 >2 > 10 possible < 500

12.2 Basic Requirements for Ceramic Powders and Binder Systems

Ceramic injection molding is a versatile shaping method, as almost every ceramic material is available in powder form. To be optimally suited for CIM, the powder should feature a number of characteristics and properties. In general, a decrease in particle size results in a higher viscosity [7] and a lower moldable solid content of the powder–polymer mixture [8]. For that reason, often relatively coarse powders with a mean particle size > 1 lm are preferred for injection molding. However, in the case of microparts, submicron powders have to be used to ensure precise replication of the microscopic details and a good surface quality, and also to achieve homogeneous properties on a micrometer scale. Fine powders also exhibit high sintering activity, allowing for a good densification and providing a high strength of the naturally fragile ceramic bodies. Ceramic powders with a mean particle size in the region of 0.5 lm are preferred for lCIM. Such powders allow the replication of details in the range of a few micrometers with sufficient quality. A general finding is that, irrespective of the mean powder size, the powder moldability decreases as the particle size distribution narrows (Fig. 12-1). Broad or multimodal size distributions are, therefore, preferred to narrow distributions to decrease the viscosity [4]. With bimodal powder mixtures, the packing density can be increased, especially when the particles are very different in size. For a high packing density, the particle shapes should be spherical or rounded or at least equiaxial. On the other hand, problems can exist for these useful powder characteristics. For example, a broad particle size distribution can result in extensive grain growth [9] and spherical particles reduce the strength of the after binder removal compacts, whereas in the presence of irregular particles the handling is eased as the green strength is higher [3]. Owing to adhesion forces between the particles, agglomeration is natural in ceramic powders. In particular, fine powders, which are obligatory for the shaping of microparts, show an increased tendency to build agglomerates because of their high specific surface area. As agglomerates may cause cracks during the debinding or sintering steps and strength-limiting defects in the final sintered

12.2 Basic Requirements for Ceramic Powders and Binder Systems Fig. 12-1 Viscosity of a dispersion of particles in a liquid decreases as the ratio R of large to small particle size increases. The figure shows R values from 1 (monomodal) to nearly infinity. Reprinted from [4], Copyright 1995, with permission from Kluwer

product [10], they must be destroyed in the feedstock compounding process and the resulting particles must be distributed homogeneously throughout the organic matrix. However, experiments showed that even with very intensive mixing of the compound not all agglomerates will be disintegrated [11]. Nevertheless, a powder that is nearly free of hard agglomerates is essential for the preparation of a feedstock for the fabrication of microparts. The lack of plasticity is characteristic of non-clay ceramic powders. However, plasticity is essential for the injection molding process, so the powders have to be blended with an organic vehicle to build a compound called the feedstock. The organic components should not only allow flowability for the filling of the mold, but also provide the green strength for defect-free ejection of the solidified body from the cavity. Furthermore, organic media are used to ensure defect-free solidification in the mold and fast and defect-free removal from the powder compact, e.g. by pyrolysis or dissolution [12]. The organic binder system usually consists of several organic components. The general range of the feedstock properties is determined by a major binder component which is normally a thermoplastic polymer or a wax-like organic. New binder systems which are based on water-soluble organics promise benefits for the debinding process. Their low green strength, however, prohibits application to fragile microparts [5]. Minor components, such as plasticizers, are added to modify the flow behavior. Dispersants improve the wetting of the particles and stabilize the dispersion against reagglomeration. The amount of binder depends on the powder characteristics and covers a range from 15 to 50 vol.% of the feedstock [13]. In general, it is desirable to attain a high solid content in the mixture while maintaining a low feedstock viscosity. A solid content well below 50 vol.% causes high shrinkage during the debinding and sintering process, leading to undesirable results such as poor dimensional control or cracking of the part.

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12.3 High-pressure Ceramic Injection Molding (HPIM) 12.3.1 Feedstocks for HPIM

Several commercial binders and ready-to-mold feedstocks are already available for the production of CIM parts. The most common binder systems are Licomont (Clariant, Gersthofen, Germany) and Siliplast (Zschimmer & Schwarz, Lahnstein, Germany). These binders are mixed with the ceramic powder to form the feedstock. Both binders are also available as ready to mold feedstocks (e.g. Inmafeed based on Licomont or Elutec based on Siliplast). Another established feedstock system is Catamold (BASF, Ludwigshafen, Germany). All feedstocks are available for many common ceramic and metal materials, designed for a broad range of applications and fitting most requirements of macroscopic fabrication. The binders Licomont and Siliplast are based on polyolefin wax and polyvinylalcohol, respectively. The concept of these binders is solvent debinding, either completely (Siliplast) or partly (Licomont), using e.g. water or alcohol. The remaining organic components necessary to maintain the stability of the debindered part, the so-called brown part, are removed during sintering. In contrast to the binders Licomont and Siliplast, the Catamold system is debindered catalytically. Thereby the main component polyoxymethylene (POM) is decomposed in a furnace using nitric acid as catalyst. Owing to the naturally higher strength of POM, the green stability of molded parts is higher than for parts made using Licomont- or Siliplast-based feedstocks. However, the Catamold feedstock shows high viscosity even at higher temperatures, making form filling of fine details difficult. Commercial binders and feedstocks are designed for general use and for the production of parts in the centimeter or millimeter range. They are optimized for short cycle times and easy post-processing, e.g. a fast debinding procedure. However, for the production of microparts, the use of these systems is limited. For microparts made by PIM, parameters such as homogeneous green density and easy defect-free demolding are crucial owing to the much higher surface-to-volume ratio. The smaller the produced parts are, the greater are the demands on tolerances and surface quality. To comply with these requirements, the use of specially designed binders and feedstocks with fine-sized powders is essential. Therefore, it is necessary to use specially designed binders consisting of appropriate thermoplastic components, waxes and additives providing, e.g., better flow behavior or higher powder–binder interaction [14]. Subsequently these binders have to be mixed together with suitable fine-sized powders to obtain a feedstock that fulfils the requirements for the production of microparts.

12.3 High-pressure Ceramic Injection Molding (HPIM)

12.3.2 Binders for MicroHPIM (lHPIM)

The basics of the binders used for the development and production of CIM feedstocks are similar to those for the binders for MIM feedstocks. For the production of microparts or micropatterned parts, they contain a significant amount of at least one thermoplastic component providing sufficient green strength for damage-free demolding of the green parts. Similarly to MIM, thermoplastics such as polyoxymethylene (POM), polyethylene (PE), polypropylene (PP), polyamide (PA) and polyvinylalcohol (PVA) or modified versions of these components are used as thermoplasts. They are often combined with waxes and further additives to form a binder system usually consisting of at least three, and often up to five, components [13]. Binders consisting of several components are preferred to realize a stepwise debinding process in order to avoid high pressure or stress in the green compacts during debinding. The waxes are used to reduce viscosity to improve the mold filling [15] and to provide better debinding behavior [8, 16]. Further details of these components are described in Section 3.2.2.1. The main difference in binder composition compared with MIM is the application of additives, so-called surfactants or dispersants, for better powder–binder interaction. Binders are based on hydrocarbons and are, therefore, non-polar. The oxide ceramic powders are polar and reactive to water, building hydroxides on the surface [17, 18]. This leads to a poor interaction between powder and binder when pure thermoplastic polymers or waxes are used. Better wetting of the polar surface of ceramic powders with the non-polar binder can be achieved by using surfactants. Relevant molecules are based on non-polar hydrocarbon chains (chain length often C16 or C18) carrying a functional polar group such as carboxylic acids (e.g. stearic acid [13, 19, 20]), amines, esters and ethers. The interaction of the hydrophilic end of a surfactant with the polar sites of the powder results in a powder surface with hydrocarbon molecules, that can be mixed better with the non-polar thermoplastic- and/or wax-based binder components leading to more homogeneous compounds. The use of additives is even more crucial when fine ceramic powders with particle size down to the submicron range are processed. Owing to the about one order of magnitude smaller particle size compared with metal powders, ceramic powders usually have a much higher specific surface area interacting with the surfactant and the other binder components during the mixing process. 12.3.3 Compounding the Feedstock

Mixing of powder and binder is an essential step in the CIM process, as the feedstock properties have a large influence on the following steps of injection molding, debinding and sintering. Especially for the shaping of microparts, the CIM feedstock must have a low viscosity at higher shear rates, even for a small powder particle size which is necessary for complete filling of the small cavity

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Fig. 12-2 Machines for feedstock preparation: kneader (W 50 EHT, Brabender, Duisburg, Germany) and extruder (ZSE 27 HP, Leistritz, Nürnberg, Germany)

details of the mold. A high powder content is preferred for the reduction of the shrinkage during sintering to increase the geometric accuracy. As with higher powder content the feedstock viscosity increases, a good CIM feedstock is always a compromise between high powder content and sufficient flowability. Besides, the compounding method has to ensure good homogeneity of the mix to avoid problems regarding injection molding or during post-processing. For mixing CIM feedstocks, there are different types of machines available. Kneaders are used for development purposes, making small batches of material, whereas twin screw extruders or shear roll compactors are continuously operating machines commonly used for commercial feedstock production on a large scale. Mixing of powder and binder is carried out at temperatures above the softening point of the binder. The lowered viscosity of the binder components allows efficient mixing with the powder (see Chapter 11.2). When fine powder particles are used, this is often accompanied by increasing powder agglomeration. Better feedstock homogeneity can be achieved when these powder agglomerates are destroyed by applying sufficiently high shear rates. Because of the serrated arrangement of the screws (Fig. 12-2), higher shear rates can be realized in an extruder, compared with mixing in a kneader where only tangential forces are acting. However, the shear rates that may be applied for the destruction of hard agglomerates have to be limited to a certain level, depending on the binder constitution, to prevent chain degradation of the binder. This can be realized with an adapted screw configuration and appropriate adjustment of the extrusion process itself regarding, e.g., rotation speed of the screws and overall machine loading. For development purposes, especially for scientific investigations, a measuring mixer with a small mixing volume is advantageous [21]. This machine allows

12.3 High-pressure Ceramic Injection Molding (HPIM)

Fig. 12-3 Torques of zirconia feedstocks with different solid loadings, measured in a kneading measuring mixer (Brabender W 50 EHT)

the measurement of material temperature and applied torque necessary to maintain the preset rotating speed of the rotors. If the loading of the machine is kept constant, the measured torque directly correlates with the viscosity of the investigated feedstock. The example presented in Fig. 12-3 shows the development of torque as a function of the solid loading. Measurement for different solid loadings was established to determine the critical loading for zirconia powders in a commercial binder (Licomont). As expected, torque (and hence viscosity) rises with increased solid loading [19]. A mixed feedstock is said to be homogeneous as soon as the measured torque reaches a constant level during mixing and when fluctuations are low, especially at the end of the mixing process. Data show that for the given example, a homogeneous feedstock is obtained for a solid loading of 56 vol.% powder or less. If this powder content is slightly exceeded, the proceeding powder deagglomeration leads to an increase of the mixing torque. A further increase in the powder content leads to strong scattering of the measured torque introduced by the remaining powder agglomerates. With respect to the production of microparts, it is necessary to limit the solid loading to a value well below the critical loading. This gives a certain reserve to guarantee automated production even when some agglomerates still remain in the feedstock after mixing. 12.3.4 Rheology

For injection molding, especially for lCIM, where molds with fine details have to be filled, the moldability and the mold filling behavior of the feedstocks used are crucial. For proper mold filling, the viscosity of a feedstock has to be sufficiently low during the injection process. To ensure this demand, the measure-

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Fig. 12-4 Viscosity of zirconia feedstocks with a Licomont binder using different zirconia powders, measured in a high-pressure capillary rheometer (Rheograph 2003, Göttfert, Buchen/Odenwald, Germany)

ment of rheological data in a high-pressure capillary rheometer (HPCR) (Rheograph 2003, Goettfert, Buchen/Odenwald, Germany) is a valuable tool for improvement of the compound quality, e.g. binder composition or powder loading. Thereby the viscosity of a compound is measured as a function of the applied shear rate at a given temperature. Regarding the production of microparts, the viscosity of the feedstock during mold filling should be as low as possible, preferably not higher than several tens of Pa · s at shear rates of about 10 000/s or more. The powders used for CIM are usually finer than those used for MIM. This is the reason why the solid loading of CIM feedstocks is lower than those of MIM feedstocks at a comparable viscosity. Furthermore, the smaller powder particles of the CIM feedstock with their higher specific surface area lead to higher powder–binder interaction. Therefore, the phenomenon of powder–binder separation, described in Chapter 11.2.5, is less critical for CIM feedstocks. In contrast to the rheological measurement of metal feedstocks, the measurement of ceramic feedstocks is possible at higher shear rates provided by nozzles of smaller diameter and/or higher travel speeds of the piston. Especially when using zirconia feedstocks with an average powder particle size of 0.35 lm and solid loadings of *50 vol.%, viscosity can be measured at shear rates of 10 000/s and more. As an example, data for feedstocks with solid loading of 48 vol.% using zirconia powders of different producers are shown in Fig. 12-4. Nevertheless, the measured apparent viscosity of the feedstocks produced differs over a wide range although the particle size and specific surface area (BET) of the different powders are similar. This indicates that particle shape and surface chemistry interacting with the binder are crucial for the viscosity of a feedstock, giving an

12.3 High-pressure Ceramic Injection Molding (HPIM)

Fig. 12-5 Viscosity of zirconia feedstocks with thermoplast/wax binder and Licomont measured in a high-pressure capillary rheometer

indication of the applicability of the respective type of powder regarding the fabrication of suitable feedstocks. Especially for the production of microparts, mold filling and damage-free demolding are decisive steps. Hence it is necessary to adapt the feedstocks to good moldability and high green strength. The characteristics of the viscosity to fit these purposes are illustrated with the data presented in Fig. 12-5. The graph demonstrates two similar zirconia feedstocks with almost identical solid loading, one with a commercial binder (Licomont) and the other with a binder based on a thermoplastic component and a wax. Comparison of the rheological data indicates that both feedstocks show shear thinning behavior. However, the viscosity of the feedstock with the commercial binder decreases less with increasing shear rate than for the feedstock with the thermoplast/wax binder. This means that the thermoplast/wax-based feedstock shows higher flowability and better form filling when injected into the mold during processing at high piston/screw traveling speed, and also better stability after mold filling due to low flowability at low shear rates. This is helpful for maintaining the shape during solidification. These rheological characteristics show that for lHPIM the feedstocks based on the thermoplast/wax binders are preferable to those with the commercial binder. In addition to the measurement of the rheological data, a high-pressure capillary rheometer is an effective tool for measuring pVT data. Thereby the dependence of pressure (p), volume (V) and temperature (T) is measured. The significant volume changes at specific melting temperatures induced by the different components of the binder give the opportunity to deduce parameters for injection molding or debinding of the green parts. Furthermore, pVT data are needed for the simulation of the microinjection molding process regarding feedstock flow characteristics and mold filling (described in Chapter 3). Fig. 12-6 shows a

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Fig. 12-6 pVT data for a zirconia feedstock with Licomont binder

pVT diagram in typical isobaric presentation for a feedstock produced with a commercial binder. The specific volume at defined temperatures measured for several pressures shows the behavior of the feedstock when pressure and temperature are applied. For the presented zirconia feedstock the investigation shows that the binder consists of at least two significant components. One melts at temperatures of about 50 8C and the other at around 140 8C at normal pressure. Therefore, the temperature for injection molding has to be at least 150 8C to ensure that the material is properly molten and the viscosity is low enough for complete mold filling. Although the calculated diagram is drawn in isobaric presentation, measurement is performed in the isothermal mode when using a high-pressure capillary rheometer, as the variation of temperature is difficult to establish. This means that the resulting volume of a sample is measured for different applied pressures at a preset temperature. Subsequently the temperature is varied and measurement of the volume as a function of the applied pressures is repeated for each temperature. 12.3.5 Molding Process

In terms of machinery, there is basically no difference between metal and ceramic lPIM. Principles, requirements and further details were given in Chapter 11. Variations in the injection molding parameters are mainly a function of the binder composition; the filler material plays only a subordinate role in spite of the different powder particle size. Differences in the flow properties of metallic and ceramic feedstocks can be balanced to a certain extent by adjusting the sol-

12.3 High-pressure Ceramic Injection Molding (HPIM)

id loading. Owing to the fact that the average particle size of ceramic powders is about one order of magnitude smaller than, e.g., steel powders, the achievable solid loading is far lower for ceramic feedstocks. As a result, a longer cooling time of the tool may be necessary for ceramic feedstocks because of the lower thermal conductivity of polymers and ceramics in comparison with metals. Typically, CIM shows more shrinkage during the sintering process, and consequently the dimensions of the molds are different to those with MIM. This means that for sintered parts with identical sizes made from metals or ceramics, two sets of tools and mold inserts have to be manufactured. Up to now, microparts have usually been injection molded on a substrate plate which is required for the demolding step. In Chapter 11.3.2, examples of microparts made by metal injection molding (MIM) are shown which can also be manufactured by CIM. The outer diameter of the smallest involute toothed gearwheel is 550 lm in the sintered state. The major disadvantage of the use of a substrate plate is that the microparts have to be isolated mechanically by grinding or mill cutting the substrate plate. This mechanical machining leads to microdefects on the surface and to residual stresses, which influence the mechanical properties of microparts significantly. By further tool developments, e.g. a three-plate mold which separates the micropart from the runner during the opening movement, this substrate plate can be omitted. Figs. 12-7–12-9 show a variety of injection molded microsized and micropatterned ZrO2 parts. The ceramic rings are injection molded in a fully automated process using a fan gate and automated separation of micropart and runner system. The diameter of the hole is 160 lm in the green state. The scale of the ruler in Fig. 12-9 is in centimeters. On the left side in the center a base plate of a planetary gear can be seen; the central planetary gear is located on the ruler in the lower right. A tensile test bar is shown on the upper border, surrounded by nozzle plates of a microturbine. The smallest channels of these nozzle plates are 25 lm in width.

Fig. 12-7 Ceramic microrings (ZrO2, green parts)

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Fig. 12-8 SEM image of a ceramic microring. Bore diameter: 160 lm

Fig. 12-9 Various sintered ZrO2 microparts made by HPIM

12.3.6 Thermal Treatment

The parameters for debinding are primarily dependent on the binder components, whereas the temperature profiles and required atmospheres and pressures for the sintering process are governed by the material. The basic princi-

12.3 High-pressure Ceramic Injection Molding (HPIM)

ples of debinding green compacts and sintering of nearly binder-free (‘brown’) compacts are given below. After shaping by injection molding, the binder has to be removed and the parts must be sintered to obtain dense compacts of the desired material. The removal of the binder is one of the most critical steps of the process, from both the technical and economic points of view. Debinding can be carried out by thermal degradation, solvent extraction or catalytic decomposition of the binder components or by a combination of these methods. Prior to the described debinding steps, a wax or an oily component of the binder can be absorbed by wicking in a powder bed. Owing to the fact that handling of microparts is fairly difficult and that the parts have to be removed from the powder bed and rinsed, the method is disadvantageous in this case. For large parts, the debinding time depends considerably on the wall thickness. Usually this limits the economics of the whole process to a wall thickness below 5 mm. However, for microparts with a wall thickness significantly below 1 mm, debinding times are less critical but still in the range 8–24 h. Here the interaction of the material with the sintering tray and the conservation of the shape are more important than shortening the debinding cycle. Furthermore, one run in the debinding furnace will produce a large number of parts, so the operational capacity is very high even with long debinding times. Exemplary temperature–time profiles for debinding and sintering of CIM parts with and without preliminary solvent extraction are shown in Figs. 12-10 and 12-11. A preliminary solvent extraction of some binder components creates an open porosity in the green compact, thus reducing the time required for the thermal debinding step and allowing higher heating rates and shorter residence times. The temperature profiles during debinding have to be controlled very accurately to avoid distortion and the formation of cracks and bubbles. When an open porosity has been achieved after the initial debinding stage, the polymeric component should still bind the powder particles together while the temperature is increasing slowly. The part should be completely free of polymer as soon as the first sinter necks have formed. Binder decomposition has to be avoided to prevent residues which influence the base materials and the microstructure

Fig. 12-10 Representative temperature– time profile for the thermal debinding and sintering of a ceramic feedstock (ZrO2)

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of the final parts. In the case of carbon removal, this process can be assisted and accelerated to a certain extent by flushing air through the debinding furnace. To avoid an additional handling step combined with cooling and heating the parts one more time, it is advantageous if the subsequent sintering process can be carried out directly after the debinding step using the same furnace. The sintering conditions, such as temperature profile and atmosphere, mainly depend on the type of material processed. Batch furnaces running under inert or oxidizing atmospheres are fairly common. More recently, semi-continuous running beam furnaces with separated compartments for debinding and sintering were introduced. Usually the parts experience a linear shrinkage of between 15 and 22% during the sintering process, depending on the powder loading, the material used and the density reached.

12.3.7 Example of a Typical lHPIM Production Cycle

The following example describes the complete production process for the manufacture of microtensile test specimens, as shown in Fig. 12-12, and bending test

Fig. 12-12 Test specimen for the determination of mechanical properties of microparts

12.3 High-pressure Ceramic Injection Molding (HPIM)

bars with a size of 0.26 ´ 0.26 ´ 3.25 mm3. For feedstock preparation, 48 vol.% of ZrO2 powder type 3YS-E from Tosoh is mixed with 52 vol.% of the thermoplast and wax binder components at 130 8C in a laboratory mixer with torque measurement (W 50 EHT, Brabender, Duisburg, Germany) for 1 h at 30 rpm. To obtain a moldable feedstock, the registered torque must not exceed 35 N m at the end of the mixing period. The rheological properties of this mixture are determined with a high-pressure capillary rheometer (Rheograph 2003, Goettfert, Buchen/Odenwald, Germany), which provides a granulated feedstock directly after carrying out the measurement. This feedstock is injection molded on a microinjection molding machine (MicroSystem 50, Battenfeld, Kottingbrunn, Austria). The temperatures of the barrel are 60–140–145–150 8C from the inlet to the nozzle. The injection speed is in the range between 200 and 400 mm/s depending on the structure of the cavity. The temperature of the tool is 60 8C when running an isothermal process. When the variotherm technique (see Section 11.3.3) is applied, the mold temperature is 85 8C for injection and 60 8C for demolding. The injection molded microparts are directly placed on sintering trays by an automatic handling system. When the sintering trays are completely filled, they are put into a closed vessel containing an organic solvent (e.g. hexane) for the first step of the debinding process where the wax component is partially extracted at 40 8C for 8–12 h. After drying the parts for 3 h at 50 8C in a vacuum drying chamber, the thermal debinding step and the sintering process are carried out in a sintering furnace (RHF-1700, CarboLite, Ubstadt-Weiher, Germany). The parameters used in this example were developed and optimized to achieve dense compacts without any cracks, bubbles or distortions and are listed in Table 12-3. The as-fired microbending test specimens are used for the determination of mechanical properties. Results of these investigations can be found in Chapter 20. The mean shrinkage during sintering is 22% linear. To achieve sufficient mechanical properties, the density should reach at least 98% of the theoretical density (6.05 g/cm3; value given by Tosoh). The manufactured microtensile test and bending test specimens show porosities between 0.3 and 0.5% and densities of 5.98–6.03 g/cm3. The measured bending strength is in the range 1400–1850 MPa, whereas the value for macroscopic specimens is 900–1000 MPa (given by the powder manufacturer). The grain sizes in the sintered bodies are well below 2 lm.

Table 12-3 Parameters for thermal debinding and sintering of ZrO2 microbending test specimens

Start temperature (8C)

End temperature (8C)

Heating rate (K/min)

Hold time (min)

50 180 270 600 1500

180 270 600 1500 100

0.2 0.5 2.0 5.0 5.0

60 30 60 60

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12.3.8 Special Development: Two-component MicroCIM

A brief introduction to two-component PIM and two-component MicroMIM was given in Chapter 11.5, describing a two-component tooling, injection molding machine and the design of a tensile test specimen with the accompanying mold insert. The described equipment and mold insert were also employed at the Forschungszentrum Karlsruhe, Germany, in work on two-component lCIM [22, 23]. This research was concentrated on the development and molding of alumina/titanium nitride feedstocks for electrically conductive ceramics. Using two different electrically conductive ceramic mixtures, a simple heater was manufactured for demonstration purposes. Apart from functioning as a heater, other fields of use could include, for example, electrically conductive components working under extreme mechanical, tribological or chemical conditions. With the aim of producing two-component ceramic parts with adjustable electrical resistivities, a mixture of insulating aluminum oxide (Al2O3) and electrically conductive titanium nitride (TiN) was chosen [24]. Because of the almost identical thermal expansion coefficients of the two materials, they are particularly suitable for this purpose. Moreover, both materials are chemically inert against one another and TiN shows metallic conductivity with a positive temperature coefficient of resistivity. It is also important that sintering can be carried out under a nitrogen atmosphere. Electrically conductive carbides, for example, have to be sintered under a noble gas atmosphere. Feedstocks were prepared containing *50–60 vol.% of powder. An exemplary feedstock composition would be one with a volume ratio of Al2O3 to TiN of 60:40 (ATN40) and 16 wt.% binder. A feedstock of this kind contains 54.8 vol% of the powder mixture, and, in particular, 21.9 vol.% TiN. Fig. 12-13 shows green and sintered two-component tensile test bars, injection molded using the mold insert shown in Fig. 11-36. The cross-section of the narrow part of the tensile test bar is 0.48 ´ 0.48 mm2 in the green part and *0.41 ´ 0.41 mm2 in the

Fig. 12-13 Two-component tensile test bars containing two different mixtures of Al2O3 and TiN (ATN). In each case the interface between the mixtures lies in the middle of the narrow section of the part. Green (top) and sintered (bottom) parts. Smallest cross-section of the green part is 0.48 ´ 0.48 mm

12.3 High-pressure Ceramic Injection Molding (HPIM)

Fig. 12-14 Two-component ATN ceramic heater. (a) green part. Dimensions: length = 22 mm, cross-section of the legs = 1.0 ´ 1.0 mm. The heating zone of the part (black) possesses a lower electrical conductivity than the legs. (b) A sintered part undergoing a functional test as a heater

Fig. 12-15 Tensile test specimen made of two different ATN ceramics. (a) Failure outside the boundary zone. (b) EDX mapping (Ti) of a boundary zone in an ATN27–ATN40 tensile test specimen. High and low TiN concentrations are clearly separated

sintered part. The interface between the two different ceramic mixtures, which contain 35 and 40 vol.% TiN, respectively, is in the center of the narrow part of the tensile test bars. Fig. 12-14 a shows a two-component heater made of an ATN feedstock (green part). The feedstock in the tip of the heater is colored with carbon black for easier detection of the materials interface. The penetration of the feedstocks into one another does not seem to be extensive. On the other hand, tests of the bonding strength in sintered parts show that they do fail outside the interface area (see Fig. 12-15 a), indicating strong interfacial bonding of the two components. Fig. 12-14 b also shows a sintered heater glowing in a functional test. An applied voltage of 4.5 V results in a current of 3.5 A, corresponding to a power uptake of around 16 W. Fig. 12-1 5b shows an EDX mapping of Ti in the bonding area. High and low TiN concentrations are in this case separated by a rela-

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tively sharp line. Despite an injection molding process which allows the feedstocks to combine in the molten state, a minimal amount of intermixing takes place. This can certainly be attributed to the high viscosities of the feedstocks employed. With regard to the research described here, a substantial amount of development work is still necessary, e.g. concerning the reproducibility in molding, debinding and sintering. Generally, research on two-component lCIM should be focused on more fundamental topics such as the adjustment of the sintering shrinkage of feedstocks, which is a challenge. The behavior of the feedstocks at the boundary between two components during injection has to be investigated in addition to the mechanical properties of two-component parts. Further miniaturization and an expansion of the materials range would also be interesting tasks.

12.4 Low-pressure Injection Molding (LPIM) 12.4.1 Characteristic Features of LPIM

Low-pressure injection molding (LPIM) is a variant of the powder injection molding process (PIM), which is virtually identical with the conventional highpressure injection molding (HPIM). However, there is one essential difference between the two processes: the pressures employed in LPIM are in the range 0.1–1 MPa, whereas HPIM takes place at pressures of > 50 MPa. This difference arises from the use of a low-viscosity paraffin or wax instead of a high-viscosity polymeric binder. Owing to the different binder systems, slight differences between HPIM and LPIM also occur for the feedstock preparation, the injection molding machinery and the debinding process. Nevertheless, the methods are closely related as the basic principles of plastic shaping are common for both processes. LPIM, which is sometimes also called hot molding [25], was invented by Gribovski in the former Soviet Union [26], where the process gained similar importance as its high-pressure counterpart in the USA, Western Europe or Japan. Recently, LPIM has been strongly propagated as a method for prototyping and small series fabrication. In contrast to HPIM, with its high costs of tooling fabrication requiring mass production for a return of investment, LPIM can work with simple and inexpensive molds and is economic even for a small number of parts. The good flowabilty of the low-viscosity feedstocks and the ability to employ fine-sized powders also recommend LPIM for the injection molding of microdimensional devices [27]. The major drawback of LPIM is the low mechanical strength of the binders used. This can lead to rupture of patterns during the demolding of the green compact, especially when fine particulars with high aspect ratio (height to

12.4 Low-pressure Injection Molding (LPIM)

width) must be withdrawn from the cavity. For that reason, LPIM is normally limited to parts with lower complexity. Although this problem can be overcome by customized tooling concepts, such as soft tooling with silicone rubber molds [28] or lost molds [29], this also prevents the automation of the process for complex parts and restricts LPIM to smaller quantities. The LPIM process can also be used for the manufacture of metal parts. However, the high density of the metal powders and the low viscosity of the feedstock give rise to distinct sedimentation of the powders in the injection units and the risk of segregation during form filling of complex geometries. For that reason, the process is more common for ceramics than for metals. 12.4.2 Feedstocks for LPIM

Feedstock properties are determined by the powder [30], by the organics added and by the mixing conditions. For LPIM, paraffin or wax is used as a binder. These short-chain organics permit the preparation of feedstocks where the viscosity is distinctly reduced, e.g. by several orders of magnitude in comparison with typical HPIM feedstocks made with long-chain thermoplastic polymers such as polyethylene or polyoxymethylene (see Section 12.3). Especially for submicron or nanoscale powder particles, the use of binders with low molecular weight is beneficial as owing to the low viscosity of the unfilled material a higher solid loading of the feedstocks is possible [8]. Binder systems based on paraffin or wax are inexpensive and non-toxic. They offer beneficial dispersion properties as they produce a liquid phase with a viscosity comparable to that of water. Typically, the melting-point of paraffin is in the range 50–70 8C, depending on the mean chain length of the organic. Since the viscosity [31] and yield point [32] of a paraffin-based feedstock with high solid content are nearly invariant in the temperature range 75–85 8C, small variations of the process temperature will not lead to different flow conditions. On the other hand, the narrow melting range and the high fluidity increase the probability of deformations in the debinding step, when the green compact softens owing to the molten binder phase. Further, the debinding behavior of paraffin and wax is uncomplicated as they decompose without boiling. Their vapor pressure is negligible under normal conditions, unlike for water, and therefore the constitution of the feedstock remains constant over the processing time. Owing to the non-polar nature of the paraffin, a surface-active dispersant is necessary to improve the wetting of the polar surface of oxide powders [33]. Non-ionic dispersants have proven to be effective additives to paraffin-based binders. The polar functional group of the dispersant molecule couples to the surface of the ceramic powder; the non-polar polymer chains extend into the paraffin and stabilize the dispersion by steric repulsion [34]. Widely used dispersants are fatty acids such as oleic and stearic acid and surfactants such as alkylphenol polyglycol ether, siloxane polyglycol ether and octadecylamine [25, 33, 34]. They differ in the kind of anchoring group towards the surface, but the number of C

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atoms in the chains often lies between 16 and 18. The amount of dispersant required depends primarily on the specific surface area of the ceramic powder, but also binder–dispersant interactions may influence the dispersant addition [35]. For most ceramic powders, a recommended value in the range 0.5–2 mg of dispersant per square meter of powder surface has been found. Special attention has to be paid to careful drying of the powder used. Adsorbed water molecules have a strong affinity to the surface of oxide powders. They cover those surface sites where dispersant molecules are supposed to adsorb physically or chemically. As a consequence, the dispersants are not effective in lowering the viscosity or stabilization of the suspension if water is present [33]. The competition between dispersant and water also plays an important role in the long-term resistance against flocculation. In spite of the non-polar character of the binder, the environmental humidity has a detrimental effect on the viscosity of the feedstock [36]. 12.4.3 Compounding the Feedstock

The achievement of a homogenous feedstock for CIM depends mainly on the strength of powder agglomerates and on the shear stress imposed in the compound during mixing. Mixing time was found to be less important [37]. In HPIM feedstocks, the shear forces necessary to disintegrate the aggregated particles can be applied by slowly rotating machines, such as R-blade kneaders or twin-screw extruders, owing to the higher viscosity range of thermoplastic bin-

Fig. 12-16 Laboratory dissolver mixer, suitable for the high-speed mixing of LPIM feedstocks. Courtesy of VMAGetzmann, Reichshof, Germany

12.4 Low-pressure Injection Molding (LPIM)

der components (see Section 12.3.3). For low-viscosity LPIM feedstocks, high shear stresses can be obtained only by a fast moving mixer. A dissolver stirrer with a vane-carrying disc is such a device which is well suited for the preparation of a small amount of feedstock (Fig. 12-16). Circumferential velocities of 10–25 m/s generate alternating stresses before and behind the vanes which are effective in wetting the particles by the liquid phase and disintegrating the agglomerates. Heated three-roller mills [31] or ball mills [25] are alternatives for the homogenization and deagglomeration of LPIM feedstocks. Usually, feedstock mixing is performed at temperatures from 80 to 100 8C for less than 1 h. To avoid trapping of air bubbles during the stirring process, evacuation of the container is advisable. However, one must be aware that reduced pressure in combination with higher temperature can lead to partial evaporation of binder and dispersant components which may be not negligible, regarding their influence on the rheological properties of the feedstock. 12.4.4 Rheology

Some requirements on the feedstock rheology are conflicting to a certain extent. Good flowability allows complete mold filling even at low molding pressures. It eases degassing and removal of bubbles from the feedstock. On the other hand, a fluid-like feedstock behavior promotes sedimentation of the powder particles during storage and processing and it complicates shape retention in the debinding step. A good compromise can be found in systems where shear thinning and a yield point exist. Unlike Newtonian behavior, where viscosity is independent of shear rate, in the case of shear thinning a viscosity decrease is observed with increasing shear rate. Especially for the injection molding of microparts this is an important feature, as small cross-sections lead to high injection speeds and thereby high shear rates. The opposite behavior – shear thickening – has to be avoided by all means, as filling of fine channels or micropatterns is nearly impossible when the viscosity of the feedstock increases with increasing shear rate. When shear thinning is accompanied by an adequate yield point, the risk of deformation of molded parts is reduced during the thermal processing. Adequate means that a minimum force is required to cause the feedstock to flow. This force has to be high enough to prevent both sedimentation of particles and deformation under the influence of gravity and surface tension. On the other hand, the yield stress should be low enough to allow the escape of enclosed air bubbles under vacuum treatment. Extensive removal of air becomes increasingly important with decreasing feature size. With respect to microparts, eventually the resulting pores are of the dimensions of the feature‘s cross section and will lead to severe defects. In Fig. 12-17, examples are shown of the shear stress and shear viscosity of various submicron zirconia feedstocks (TZ-3YS-E, Tosoh Europe, Amsterdam,

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Fig. 12-17 Influence of different dispersants on the rheological properties of paraffin-based zirconia feedstocks at 85 8C. (a) Shear stress as a function of shear rate; (b) shear viscosity as a function of shear rate

The Netherlands) with dispersant systems added to the paraffin on-site and with a commercially premixed binder system. Shear-thinning behavior is observed for all three feedstocks with decreasing shear viscosity with increasing shear stress in Fig. 12-17b. By using various additives, the yield point and flow behavior at low shear rates can be adjusted to the desired values. To compare the rheological properties of different feedstocks, the viscosity is often selected at a shear rate of 100 s–1. Adequate moldability for LPIM was found with viscosities of < 15–20 Pa s at 100 s–1.

12.4.5 Machines, Tooling

In principle, a standard injection molding machine, which is used for HPIM, can also be charged with a LPIM feedstock. However, these machines are designed for a high-pressure regime and do not operate properly in the low-pressure range. Under high-pressure conditions, the benefits of the low pressure molding, e.g. inexpensive tooling, are lost. Additionally, at high pressure or high flow rates, a controlled mold filling is difficult, as jetting, splashing or air entrapment can take place [12]. For that reason, special LPIM machines were developed that can work at a molding pressure of < 1 MPa. Two types of LPIM machines can be distinguished: pneumatically driven machines and piston-type machines. Examples of the first category are the MIGL machines (Fig. 12-18 a and b), manufactured by Peltsman (Minneapolis, MN, USA). Here the feedstock is pressed into the mold by compressed air. A built-in mixer prevents the segregation or sedimentation of low-viscosity feedstocks. Owing to the simple concept, home-made constructions with a pneumatic injection principle also exist in some laboratories and factories.

12.4 Low-pressure Injection Molding (LPIM)

Fig. 12-18 Low-pressure injection molding machines. (a) Schematic diagram and (b) photograph of an MIGL-33 from Peltsman. Courtesy of Peltsman, Minneapolis, MN, USA. (c) Photograph of a GOCERAM GC-MPIM-2-MA. Courtesy of GOCERAM, Mölndal, Sweden

A piston-type machine is fabricated by GOCERAM (Mölndal, Sweden). The use of a piston allows molding pressures of up to 7 MPa, but the machine is also suitable for the low-pressure regime. In the standard machine, the feedstock is injected into the mold by a pneumatically driven piston (Fig. 12-18 c). Another model is equipped with an electrical servo motor to run injections with different rate profiles. As for HPIM, stainless steel is the preferred material for standard LPIM tools, but the tooling for LPIM can have a much lower complexity and requires less cost and time for manufacturing. Molding conditions are moderate compared with HPIM (see Section 12.3.7). Typical feedstock temperatures are in the range 70–90 8C at an injection pressure below 1 MPa. Owing to the good flowability of

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the feedstocks and low molding pressures and temperatures, there is no need for a massive clamping unit or for variotherm process features (described in Chapter 10). In addition, LPIM feedstocks generate less wear. Hence LPIM works also well with tools which are made of easily machinable materials such as brass, aluminum and polymers. As this allows the fabrication of inexpensive molds, LPIM has benefits for the production of prototypes or small lot sizes. Silicone rubber is another material which can be used for LPIM tools (soft tooling). In spite of the softness of silicone rubber, it is possible to prevent distortion of the mold at a sufficiently low molding pressure (usually l < 0.3 MPa). This has a tremendous benefit for the demolding of micropatterned parts [38]. Owing to the elasticity of the material, it is possible to withdraw parts with parallel side walls or even undercuts from the cavity, whereas in non-compliant materials a demolding of the structure is challenging or hardly possible. Some examples of ceramic microparts which have been manufactured by LPIM from silicone rubber molds are presented in the following chapters. Another option of LPIM is the use of lost mold materials, i.e. molds made of wax or polymer that are removed from the green compacts by pyrolysis or with solvents. Although a reusable mold seems to be less expensive than a lost mold, for the latter there is no need for elaborate ejector fixtures. This confers a cost advantage on lost molds for parts with complex shape, fragile details and high aspect ratio structures in small and medium quantities. 12.4.6 Thermal Treatment

The conversion of the green compacts to dense ceramic parts comprises two subsequent process steps. Debinding, the removal of the organic constituents, takes place at temperatures below 600 8C. For the sintering step, significantly higher temperatures are required for densification. Usually, binder removal is the most critical and time-consuming procedure in the CIM process and this step is a frequent source of defects in the ceramic product. Depending on the solid content of the feedstock, 25–55 vol.% of waxbased binder must be removed from the green compact. A considerable fraction of the binder decomposes by pyrolysis with gaseous products producing an overpressure inside of the bulk. With increasing wall thickness and with decreasing particle size, this overpressure becomes more critical and the risk of the formation of defects increases [39]. Hence the small dimensions of microparts promote the debinding of the green compact. On the other hand, the continued trend to use finer powders still aggravates the formation of critical conditions. However, whereas debinding times from days to more than 1 week have been reported for macroscopic parts [4], typically an overnight debinding is possible for microparts. Some debinding challenges that are distinctive for LPIM appear soon after exceeding the melting-point of the binder. When a liquid binder phase is formed, the green compact becomes weak and soft. Deformation of the part can occur if

12.4 Low-pressure Injection Molding (LPIM)

Fig. 12-19 Sintered ZrO2 microcolumns made by LPIM: (a) sharp edges with optimized debinding conditions; (b) rounding of edges by surface tension

a sufficiently high yield point does not exist. Whereas gravitational forces play a minor role for microparts, the surface tension of the softened body leads to plastic deformation which alters the shape of fine details by rounding initially sharp edges (Fig. 12-19) or by leveling surface patterns. Although this effect has a negative impact on the shape of the component, it can also be used for an improvement of the mechanical properties as smoothing and edge rounding also eliminate surface defects. On ZrO2 microbars with square cross-section (*200 ´ 200 lm2) and rounded edges, a tensile strength of > 3000 MPa was measured [40]. An extensive discussion of this will be given in Chapter 20. At elevated temperatures, a liquid binder phase accumulates at the surface of the green body. To sustain dimensional integrity, the emerging liquid has to be withdrawn, for example by the capillary forces of a powder bed or an absorptive support. For macroscopic parts, often a wicking powder bed is used as a mechanical support to prevent collapse and to remove the excess binder by capillary forces. However, using a powder bed has the drawback of producing a higher surface roughness, as the powder bed is in contact with the molten binder phase, producing impacts and adhering particles. In Fig. 12-20 this effect is illustrated on the basis of two ZrO2 samples differing only in the employment of a powder bed for the specimen shown in Fig. 12-20b. For large components a more rough-textured surface may not affect their applicability or, at least, this can be improved by surface finishing. However, this option is not available for microparts which have to be applied as-sintered. For that reason, a highly absorptive supporting area is essential for the debinding of microparts. In the first stage of the debinding process, open-pore channels must be created to form an interconnected network, before decomposition of the binder starts at increased temperatures and volatile products evolve. In Fig. 12-21, the representative thermogravimetric analysis (TGA) of two zirconia feedstocks with different binder–dispersant combinations is presented. Although significant

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Fig. 12-20 Sintered ZrO2 microcolumns made by LPIM: (a) debinding without powder bed; (b) debinding in a powder bed

Fig. 12-21 TGA of the debinding process for paraffin-based zirconia feedstocks (solid content 50 and 52 vol.%); mass loss as a function of temperature, heating rate 1.0 K/min, synthetic air atmosphere

mass loss does not start below temperatures of about 200 8C, volatilization of the Feedstock B takes place even at 100 8C. The highest rate is observed between 250 and 400 8C. In this temperature interval, a very slow heating rate is recommended. At 400 8C almost no mass loss can be observed, but the binder is not completely transferred into the gas phase. Reactions between binder components and their oxidation products lead to cross-linked carbonaceous residues, preferentially located at the contact points of the powder particles [41]. This residue is beneficial as it provides a minimum stability which improves the handling of the parts after binder removal. During the standard debinding process,

12.4 Low-pressure Injection Molding (LPIM)

the maximum temperature generally is not high enough to burn out the residual carbon; it is reserved for the subsequent sintering process. In addition to the classical thermal debinding, removal of the binder by solvents or by supercritical extraction is possible but will not be discussed in this chapter. Although these methods offer improved mechanical properties [42], they require further equipment and exact process control and are therefore less familiar than thermal processes. 12.4.7 Examples

Current rapid prototyping methods are characterized by restrictions on resolution and variety of materials. Especially in the micrometer range there is still a lack of suitable methods for the rapid manufacture of ceramic prototypes. A Rapid Prototyping Process Chain (RPPC), combining the fabrication of a polymer master model by rapid prototyping methods such as microstereolithography and the replication of this model into the ceramic part by LPIM using sili-

Fig. 12-22 Components of a ZrO2 microturbine. Parts were made by LPIM in silicone rubber molds (original sized parts compared with a pencil tip)

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Fig. 12-23 Microsized gear wheels, made from AlN. Reprinted from [42], Copyright 2000, with permission from DVS-Verlag

Fig. 12-24 Modular ceramic microreactor system. Individual parts are made from alumina feedstocks. From [45], reproduced by permission of the Royal Society of Chemistry

cone molds, allows this gap to be closed [28, 43]. With this process chain it is possible to produce ceramic microparts with full functionality from single prototypes to preliminary or small lot series cost-effectively and in a short time. At the Forschungszentrum Karlsruhe, RPPC was developed for the manufacture of ZrO2 microparts, e.g. elements of a ceramic microturbine (Fig. 12-22). A similar process was used at the IKTS Dresden for the manufacture of microgear wheels (Fig. 12-23) [44]. The benefits of LPIM can also be illustrated by a ceramic microreactor system (Fig. 12-24), which was developed und manufactured with an RPPC within a reasonable period of time and at acceptable cost [46]. Whereas metal microreactors had already proven to work successfully, comparable ceramic systems for very high temperatures or corrosive conditions were still lacking. For the design development and owing to the desired modular character of the reactor, a variety of molding tools was needed. As the final design could not be verified on

12.6 References

design models, but only under operating conditions, the manufacture of functional models was indispensable. With LPIM a fabrication technique was chosen that met the requirements for the molding of the relatively large reactor housings and for the micropatterned details of the modular components.

12.5 Conclusions

CIM is a versatile shaping method for the manufacture of ceramic microparts. The variants LPIM and HPIM permit the economic production of a broad spectrum of items, from a few prototypes up to full-production runs. Progress in the improvement of feedstock properties and development of adapted tooling were the preconditions for a technique which is now able to manufacture structural details down to the tens of micrometers range. The multitude of the CIM process steps, comprising feedstock compounding, injection molding, demolding, debinding and sintering, makes it necessary to examine the influence of a correspondingly high number of process parameters on important properties of the resulting microparts. Not before these relationships are completely understood can the whole process, involving both HPIM and LPIM, be optimized. An important issue is meeting the tight tolerances of only a few micrometers to maintain a high level of precision and quality, which is essential in microsystems technology. Regarding economic aspects, the tooling concepts for the production of microparts have to be improved to realize further reductions in finishing.

12.6 References 1 T. R. Haglund, US Patent 2 487 861, 2 3

4

5 6

7 8

1936. K. Schwartzwalder, US Patent 2 122 960, 1938. B. C. Mutsuddy, R. G. Ford, Ceramic Injection Molding; London: Chapmann and Hall, 1995. M. A. Janney, in: Ceramic Processing, R. A. Terpstra et al. (eds.); London: Chapman and Hall, 1995, Chapter 6. Z. S. Rak, Cfi/Ber. DKG 1998, 9, 19–26. W. Bauer, R. Knitter, V. Piotter, R. Ruprecht, Nachr. Forschungszentrum Karlsruhe 2002, 34, 229–235 (in German). R. N. Weltman, H. Green, J. Appl. Phys. 1943, 14, 569–576. J. H. Song, J. R. G. Evans, J. Rheol. 1996, 40, 131–152.

9 F. Thümmler, R. Oberacker, An Introduc-

10 11 12 13

14

15

tion to Powder Metallurgy; London: Institute of Materials, 1993. F. F. Lange, M. Metcalf, J. Am. Ceram. Soc. 1983, 66, 398–406. M. J. Edirisinghe, J. R. G. Evans, Proc. Br. Ceram. Soc. 1986, 38, 67–80. R. M. German, K. F. Hens, Ceram. Bull. 1991, 70, 1294–1302. R. M. German, Powder Injection Molding; Princeton, NJ: Metal Powder Industries Federation, 1990. S. Rath, L. Merz, K. Plewa, P. Holzer, T. Gietzelt, J. Hausselt, Adv. Eng. Mater. 2005, in print. J. H. Song, J. R. G. Evans, Ceram. Int. 1995, 21, 325–333.

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17 18 19

20 21

22

23

24

25 26 27

28 29

Streicher, C. I. Chung, in: Advances in Powder Metallurgy, Vol. 2: Powder Injection Moulding; Princeton, NJ: Metal Powder Industries Federation 1991, pp. 43–58. H. P. Boehm, Angew. Chem. 1966, 78, 617–628 (in German). J. B. Peri, J. Phys. Chem. 1965, 69, 220– 230. V. M. B. Moloney, D. Parris, M. J. Edirisinghe, J. Am. Ceram. Soc. 1995, 78, 3225–3232. A. Smolders, J. Sleurs, J. Eur. Ceram. Soc. 1997, 17, 171–175. L. Merz, S. Rath, P. Piotter, R. Ruprecht, J. Haußelt, Microsyst. Technol. 2004, 10, 202–204. G. Örlygsson, V. Piotter, G. Finnah, R. Ruprecht, J. Hausselt, in: Euro PM2003 Conference Proceedings; Shrewsbury: European Powder Metallurgy Association, 2003, pp. 49–154. G. Örlygsson, G. Finnah, V. Piotter, U. Kaufmann, R. Ruprecht, J. Hausselt, in: Advances in Powder Metallurgy and Particulate Materials – 2004; Princeton, NJ: Metal Powder Industries Federation, in press. V. Winter, R. Knitter, in: Proceedings of the MicroMaterials Conference, Berlin; 1997, pp. 1015–1017. R. Lenk, Cfi/Ber. DKG 1995, 72, 636– 639. P. O. Gribovski, Hot Molding of Ceramic Parts; Moscow: GEI, 1961 (in Russian). R. Knitter, W. Bauer, D. Göhring, J. Hausselt, Adv. Eng. Mater. 2001, 3, 49– 54. W. Bauer, R. Knitter, J. Mater. Sci. 2002, 37, 3127–3140. R. Elsebrock, C. Makovica, P. Meuffels, R. Waser, Mater. Lett. 2004, 58, 3348– 3349.

30 W. Bauer, G. Bartelt, J. Hausselt, H.-J.

31

32 33

34 35 36 37 38

39 40

41 42 43 44 45 46

Ritzhaupt-Kleissl, Key Eng. Mater. 2001, 206–213, 329–332. M. Leverkoehne. J. Coronel-Hernandez, R. Dirscherl, I. Gorlov, R. Janssen, N. Claussen, Adv. Eng. Mater. 2001, 3, 995– 998. D.-M. Liu, W.-J. Tseng, Mater. Sci. Eng. A 1998, 254, 136–146. S. Novak, K. Vidovic, M. Sajko, T. Kosmac, J. Eur. Ceram. Soc. 1997, 17, 217– 223. D.-M. Liu, J. Am. Ceram. Soc. 1999, 82, 1162–1168. S. Zürcher, T. Graule, J. Eur. Ceram. Soc. 2005, 25, 863–873. S. Novak, A. Dakskobler, V. Ribitsch, J. Eur. Ceram. Soc. 2000, 20, 2175–2181. J. H. Song, J. R. G. Evans, J. Eur. Ceram. Soc. 1993, 12, 467–478. W. Bauer, R. Knitter, G. Bartelt, A. Emde, D. Göhring, E. Hansjosten, Microsyst. Technol. 2002, 9, 81–86. J. R. G. Evans, M.aJ. Edirisinghe, J. Mater. Sci. 1991, 26, 2081–2088. D. Gronych, M. Auhorn, T. Beck, V. Schulze, D. Löhe, Z. Metallkd. 2004, 95, 551–558. J. E. Zorzi, C. A. Perottoni, J. A. H. da Jornada, J. Mater. Sci. 2002, 37, 1801–1807. T. Chartier, M. Ferrato, J.-F. Baumard, J. Eur. Ceram. Soc. 1995, 15, 899–903. R. Knitter, W. Bauer, D. Göhring, J. Mech. Eng. Sci. C1 2003, 217, 41–51. B. Alm, R. Lenk, Keram. Z. 2000, 52, 312–315. R. Knitter, M. Liauw, Lab Chip 2004, 4, 378–383. R. Knitter, D. Göhring P. Risthaus, J. Hausselt, Microsyst. Technol. 2001, 7, 85– 90.

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13 Microcasting G. Baumeister, J. Haußelt, S. Rath, R. Ruprecht, Institute for Materials Research III (IMF III), Forschungszentrum Karlsruhe, Germany Abstract Microcasting is a metal forming process based on the well-known lost-wax – lost-mold technology of investment casting. The further development of this technique for casting structures in the range of some tens of micrometers requires special patterns, investments and casting parameters. First, this chapter describes the general casting process, highlighting differences from conventional dental and jewelry casting. Additionally, the parameters of a typical microcasting process are given. Next, some alloys used for microcasting and their chemical compositions, melting and casting temperatures and phase transitions during the solidification process are described in detail. Thereafter, the two basic investments used for microcasting and the influence of the investment on the surface roughness of the cast parts are discussed. Finally, cast microparts are shown and their properties such as microstructure, dimensional accuracy, surface roughness, mechanical properties, smallest achievable structure size and highest obtainable flow length and aspect ratio are presented.

Keywords investment casting; dental casting; gold base alloy; bronze; CoCrMo alloy

13.1 13.2 13.2.1 13.2.1.1 13.2.1.2 13.2.1.3 13.2.1.4 13.2.1.5 13.2.2 13.2.3

Introduction 358 Investment Casting

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General Process 360 Process Description 360 Pattern Design 362 Melting 363 Casting 364 Solidification 365 Vacuum Pressure Casting Centrifugal Casting 367

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13.2.4 13.3 13.3.1 13.3.2 13.3.3 13.3.4 13.4 13.4.1 13.4.2 13.4.3 13.4.4 13.4.4.1 13.4.4.2 13.4.4.3 13.5 13.5.1 13.5.2 13.5.3 13.5.4 13.5.5 13.5.6 13.6 13.7

Example of a Typical Casting Process 368 Casting Alloys

369

Introduction 369 Gold Base Alloy as an Example of Precious Metals 369 Bronze as an Example of Typical Casting Alloys 371 CoCrMo Alloy as an Example of High Strength Materials 372 Investment Materials

373

Introduction 373 Phosphate Bonded Investments 375 Plaster Bonded Investments 377 Influence of the Investment on the Surface Roughness 378 Coating the Pattern 379 Infiltrating the Mold 379 Modifying the Investment 380 Cast Microparts and Their Properties

381

Examples of Cast Microparts 381 Microstructure/Grain Size 383 Dimensional Accuracy 385 Surface Roughness 386 Mechanical Properties 387 Achievable Structure Size, Flow Length and Aspect Ratio

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Conclusions 390 References 391

13.1 Introduction

Microcasting is the manufacturing process of small structures in the micrometer range or of larger parts carrying microstructures by using a metal melt which is cast into a microstructured mold. Fields of application are, e.g., instruments for minimal invasive surgery, dental devices and instruments for biotechnology. Additionally, the manufacturing of miniaturized devices for mechanical engineering is a desired outcome. At present, two different techniques for casting structures in the micrometer range are known: capillary action microcasting and microcasting based on investment casting. The first manufacturing method was developed by Bach et al. [1] and Moehwald et al. [2]. They applied capillary action microcasting for form filling of structures in the range of some micrometers. Similar to die casting, this technique uses a permanent mold which can be opened in order to remove the cast structure. The cavities in the mold are shaped by high-precision grinding [2]. For casting, two different principles to fill these cavities exist: the suction principle and the displacement principle [1]. In the first case the melt is sucked into a specially coated mold by the capillary pressure. In the second case the casting alloy is melted inside the divisible mold and fills the microstruc-

13.2 Investment Casting

tured cavities owing to the capillary force. Subsequently pressure is applied to the mold to displace the excess melt through the slit. Owing to absorption of the coating during solidification, the casting detaches from the mold’s surface, but at the same time the alloy composition changes slightly compared with the original material. In capillary action microcasting the castable geometries are limited to structures which can be filled by application of capillary forces. Microcasting based on the investment casting technique, which will be discussed in the following, does not suffer from these limitations.

13.2 Investment Casting

Microcasting, also named microprecision casting [3, 4], is generally identified with the investment casting process, a casting technology also known as the lost-wax, lost-mold technique [5, 6]. This forming process excels in near net shape manufacturing and is an established technology with great freedom in design [7]. It offers the chance to produce very complicated formed parts in metal even with undercuts. Another advantage of the investment casting process over other shaping processes is the rapidity of the casting procedure itself and the low loss of material due to the possibility of recycling the runners and sprues. However, the process cannot be fully automated, so it is best suited for small and medium series and for parts with highly complex shape. This is the reason why investment casting has, in addition to technical application, a high relevance for jewelry and dental casting. For both applications, precise manufacturing is achieved [8–13], especially by using precious alloys. For jewelry and dental casting, the sizes of the produced parts are in the millimeter up to the centimeter range with structural details in the millimeter and submillimeter ranges [1]. Further development and improvement of these techniques allowed the casting of microparts with structural details even in the micrometer range, which was confirmed by the replication of small-scale LIGA structures (see Section 13.2.1.1) with high accuracy [15]. The new microtechnology, derived from the conventional production process, requires different pattern materials, other investments, special alloys and other casting parameters compared with the standard investment casting process. Additionally, microcast parts cannot be machined mechanically after manufacturing. Sand blasting, as applied in dental and jewelry casting, cannot be used to remove residue of the investment, nor can surfaces be polished to increase their quality. Sand blasting would reduce the sharpness of the edges and therefore influence the accuracy of the part, and polishing is not possible because of their small size. Therefore, precious alloys are particularly suitable for microcasting, because here the investment can be removed chemically from the cast metal part using hydrofluoric acid without influencing the cast part. Recent progress in the development of investments, however, opened the possibility of casting microstructures with bronze as a non-precious alloy [17]. In the following, the typical microcasting process will

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be illustrated first. Later, details on the alloys and investments and variables specifically influencing this process are given. 13.2.1 General Process

The microcasting process, which is described in Section 13.2.1.1, has enormous potential in manufacturing microparts of high quality without the need for further processing, as opposed to the dental and jewelry casting technique. The patterns used in the microcasting process (see Section 13.2.1.2) guarantee a higher strength and are thus of advantage when assembling microstructures. Sections 13.2.1.3–13.2.1.5 give the basics on melting, casting and solidification.

13.2.1.1 Process Description The microcasting process itself is based on the lost-wax, lost-mold technique. It is widely comparable to casting of dental protheses or jewelry [18]. In contrast to the wax patterns used there, microtechnology mostly works with injectionmolded plastic patterns which have much higher mechanical strength. The improved mechanical properties permit easier handling and assembling of the pattern during the manufacturing process. The shaping of the microcavities in the mold insert, used for injection molding, can be achieved by several methods. We applied mainly the technique of micromilling [19–21], which is a further development and improvement of the standard milling process towards miniaturized manufacturing (see Chapter 4), but in some cases also microelectro discharge machining [22, 23]. More details on the latter process can be found in Chapter 7. Other ways for the production of microstructured mold inserts are the laser technique [24–26] and the LIGA process [28–30]. The LIGA process, which is described in detail in Chapter 8, includes a lithographic and a galvanic process and is beneficial for microreplication owing to the very good surface quality of the mold inserts and the high potential of generating minimum structures. However, in contrast to the milling process, which allows the production of free form faces and real 3D structures, the LIGA technique is limited to 2.5-dimensional structures because of their necessarily vertical walls. The microcasting process requires a lost plastic pattern to be mounted on a gate and feeding system made of wax (Fig. 13-1). The assembly is then completely embedded in a ceramic slurry. This process differs from the technical investment casting process where normally a ceramic shell is built-up by repeatedly dipping the pattern in a ceramic slurry followed by stuccoing. After drying, the ceramic is sintered, resulting in a ceramic mold with high mechanical strength. Simultaneously, the plastic melts during the burning process and is pyrolyzed. In order to fill the mold with the metallic melt, either the vacuum pressure casting or the centrifugal casting technique can be used. In the first case, the

13.2 Investment Casting

Fig. 13-1 Scheme of investment casting process

ceramic investment mold is evacuated, then the melt is poured into the mold, filling the cavity only due to gravitational forces. After that, pressure is applied to the melt. In the second case, the centrifugal force is used for form filling. Both techniques will be explained later in detail. After solidification, the investment is mechanically removed without destroying or influencing the cast surface. Depending on the casting alloy and the investment material, additional chemical cleaning processes may be sometimes necessary. Finally, the single parts are separated from the runner system. Unlike dental or jewelry casting, there is no further treatment such as grinding or polishing of the cast surface. This is due to the much smaller geometry and inapproachability of details on cast microparts and also to the necessity for high contour precision without any rounding of the edges. Fig. 13-2 shows the most important replication steps for the example of a microturbine plate. On the left, the mold insert for injection molding of the pattern, the negative form, is shown. It is made by micromilling in brass. In the

Fig. 13-2 a) Mold insert for injection molding, b) PMMA pattern, c) cast part made in gold base alloy Stabilor® G

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center is the injection-molded plastic turbine plate made of PMMA [poly(methyl methacrylate)]. This is the positive pattern required for microcasting. The plastic pattern is replicated by the investment and forms the negative mold. The third replication process – the real casting – yields the desired positive form, in this case made of a gold base alloy (right). It is worth mentioning that the replication is so precise that even scratches with depths of a few micrometers in the mold insert for injection molding are perfectly replicated on the cast part. The investment casting procedure for manufacturing microparts is influenced by many different parameters. The most important ones are the casting alloy, the ceramic investment, the preheating temperature of the mold and the casting pressure. The molten casting alloys must exhibit a low viscosity in order to fill the small microstructures completely. Additionally, a minor tendency for oxidation is of high interest. For the ceramic investment, the most important factors are the ability for high-precision replication, an expansion behavior adjusted to the alloy used and a low surface roughness. The preheating temperature and the pressure influence the entire form filling process [31] and, as a consequence, the achievable grain size and the resulting mechanical properties.

13.2.1.2 Pattern Design For cost-effective casting, the assembly of single patterns in form of a so-called tree is necessary, whereas the design rules of good castability should be considered to allow homogeneous form filling of all mounted structures. In microcasting, single polymer patterns are normally fixed with wax. As an example, Fig. 13-3 a shows a pattern with 15 injection-molded polymer tensile test specimens fixed on a sprue system made of wax. In Fig. 13-3 b, the resulting cast part (gold base alloy) can be seen. Single microstructured patterns should be made at least with a small runner owing to the difficult handling of the small parts. Forming of complete plastic or wax assemblies is even better. Especially patterns which are injection-molded on a substrate plate proved to be advantageous because the substrate plate can be used as feeder. However, the melt flow

Fig. 13-3 a) Pattern with 15 injection-molded polymer tensile test specimens fixed on a sprue system made of wax; b) equivalent cast part manufactured in a gold base alloy

13.2 Investment Casting

in the plate is not easy to control. In industry, similar problems are solved by simulating the casting process. For microdimensions such specialized tools are not yet available. Like patterns for macrocasting, patterns for microcasting should be constructed according to the well-known design rules for casting [32–35]. In order to produce faultless patterns, different wall thicknesses and sharp edges should be avoided [36]. Furthermore, the form filling process is of great importance. The cross-sectional thickness of the sprue system should increase in the direction of the sprue bottom, because solidification must begin in the microparts and end in the bottom of the tree. On the one hand, this design is beneficial for good form filling; on the other, it helps to avoid shrinkage holes in the casting. This design rule, however, does not normally cause any problems in microcasting, because the parts are generally distinctly smaller than the feeder and runner system. Nonetheless, the heat capacity of the mold should also be taken into account because the compact molds used for microcasting show a comparatively high heat capacity. This results in the inner part of the massive form being still hot while the surface cools rapidly after the mold has been taken out of the furnace. Therefore, thin-walled parts should be positioned in the outer and thick-walled parts in the inner area of the mold. The melt will then remain liquid in the thick-walled parts for the longest time so that they can work as feeder for the thinner parts. As mentioned before, an adequate sprue system is necessary in order to avoid shrinkage holes in the thick-walled parts. More detailed information on the sprue design is given in the literature [37]. A special aspect in microcasting is the flow behavior in very fine channels. Owing to the much higher surface to volume ratio in microchannels compared with macrostructures and the distinct influence of surface roughness, the occurrence of turbulent flow needs to be taken into account. Another aspect is the extremely high cooling rate and therefore extremely fast solidification in the small structures, which hinders form filling much more than in macrostructures. This aspect will be discussed in more detail in Section 13.5.6.

13.2.1.3 Melting For casting in different atmospheres, various set-ups are available. Some casting machines work with vacuum, some with air and others with an inert gas atmosphere. Also the furnaces can vary. There is electrical resistant heating, heating by an open flame, induction heating and melting by an arc furnace. For resistance heating, the heat is produced by a heat winding which encloses the crucible. The heat winding can be made, for example, of platinum–rhodium. The method is used for alloys with casting temperatures up to about 1300 8C. Such set-ups with resistance heating are predominantly used for casting precious metals. For the open flame technique, the metal is melted by a propane–oxygen flame in a ceramic crucible. The method is limited to relatively small amounts of metal and is especially used in dental casting workshops for heating high-melting

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alloys to temperatures between 1300 and 1500 8C. Here, the use of a reducing flame is of great importance in order to eliminate oxides in the melt. The open flame technique requires good craftsmanship, but given this, good casting results can be achieved. Induction heated casting machines allow for a higher automation level. They are now commonly used [38] in dental and jewelry casting because they exploit the widest range of casting atmospheres and temperatures. In induction heated equipment the metal is melted in a crucible surrounded by a water-cooled copper coil. An alternating current excites a magnetic field inducing eddy currents in the metal charge. This results in strong Joule heating due to the resistance of the charge carriers in the metal. The amount of energy injected depends on the alloy and the frequency (of the furnace). Modern equipment works with high frequencies in the region of 100 kHz. A benefit of this method is the very high melting rate. Owing to the direct injection of energy, higher melting alloys requiring casting temperatures above 1300 8C can be cast, compared with electrical resistance furnaces which are limited to 1300 8C in general. The induced eddy currents result in a strong convection in the melt. Hence good mixing and homogenization are achieved. For melting with an arc flame, a pure argon atmosphere is necessary because the gas atoms work as charge carriers for the current flow. For the same reason, the metal to be molten must have electric contact with the crucible, which is normally connected as anode. The arc is then ignited between a tungsten cathode above the crucible and the crucible itself. Arc furnaces are very powerful and are also able to melt higher melting metals such as CoCrMo alloys or titanium. The melting crucibles are made of ceramic or graphite. At high temperatures the graphite crucibles produce a reducing CO atmosphere as a result of the reaction of the carbon with the oxygen in the air. This is especially beneficial for precious metals because the melt is protected against oxidation. On the other hand, graphite crucibles tend to react with the melt so that for carbon-sensitive alloys ceramic crucibles are used normally. For titanium alloys, however, graphite crucibles are used although titanium is known to be a strong carbide former. In this case, a thin titanium carbide layer is formed in the crucible during the first melting process, which protects the melt during the following uses against reaction with the graphite crucible.

13.2.1.4 Casting Metals to be processed by microcasting must have sufficient castability. The term embraces properties such as flowability and form filling ability, little contraction and shrinkage, reduced segregation, low porosity and shrinkage cavitation, little hot crack susceptibility, high surface quality and good mechanical properties. A metal is considered castable if the mentioned properties can be sufficiently achieved by using a given casting method. Based on the criteria form filling, surface quality, microstructure and dimensional accuracy, the cast-

13.2 Investment Casting

ing quality can be judged. Deviations from the norm are regarded as casting defects. They originate either in faulty workmanship, in the selection of the wrong casting parameters or in limitations of the process. A typical problem is the solidification of the melt before the form is filled completely, which is a result of too low mold temperatures or insufficient overheating of the melt above the liquidus temperature. Additionally, incomplete casting can be caused by a filling pressure too low to overcome the surface tension of the melt, which is then unable to enter a cavity. Other important casting defects are described below. Shrinkage holes are a result of too fast solidification without sufficient feeding. Furthermore, so-called casting pearls may occur. These are metal pearls located on the casting due to primary air bubbles at the surface of the pattern which were not removed during embedding. Finally, surface shrinkage holes and surface pores caused by a too high casting temperature can sometimes be found in cast parts. Detailed information on casting defects and also images illustrating them can be found in the literature [39–42].

13.2.1.5 Solidification The molten metal is poured into a preheated form distinctly cooler than the melt. The solidification starts with nucleation and crystal growth at the cooler mold wall [43]. At the same time, the volume of the melt decreases owing to the normal shrinkage process, which may then cause casting defects [39]. Therefore, it is important that the cast part solidifies first while the metal in the sprue still remains liquid. Another important aspect is the changing of the chemical composition during the solidification due to segregation. This segregation can occur in the center of cast blocks because companion elements and inclusions are pushed aside by the solidification front and accumulate in the rest of the melt. Graduated microsegregation inside dendrites or in general inside one phase, formed during solidification, also known as coring [43], is found in alloys which show a solidification interval. In this case, a difference in alloy composition between the center and the extremities of dendrite arms occurs owing to an enrichment of one element in the forming crystals at the expense of an impoverishment of the same element in the liquid. Alloys are prone to coring if the solidification is too fast to reach an equilibrium state according to the phase diagram [44]. The chemical composition can be homogenized by a subsequent long heat treatment at relatively high temperature. 13.2.2 Vacuum Pressure Casting

For vacuum pressure casting of microparts, dental casting machines can be used. Fig. 13-4 shows a scheme of the process. The metal is melted in the crucible located in the center of a heating winding. On top of the crucible the open mold is fixed upside down. After evacuation, the machine turns itself upside down (Fig. 13-4, right). As a result, the melt flows into the mold by gravity.

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Fig. 13-4 Scheme of the vacuum pressure casting process. Left: vacuum condition with mold atop the heating chamber; right: pressure condition with melt discharged into the mold by force of gravity (machine turned upside down)

Fig. 13-5 Vacuum pressure casting machine Prestomat® by Degussa Dental GmbH

Complete form filling even of small cavities is achieved by subsequent application of pressure to the melt. Fig. 13-5 shows a view in the opened vacuum pressure casting machine Prestomat® from Degussa Dental GmbH. Vacuum pressure casting machines typically work at pressures of 3.5–4 bar, which is sufficient for the form filling of parts in the millimeter range. Depending on the special geometry of the microparts and the feeding system, this pressure may even be adequate for the casting of structures in the micrometer range. However, if extremely high aspect ratios are to be cast, a higher pressure is necessary. Calculations of the form filling behavior [45] show that the pressure which is necessary for the melt to enter an extremely small pinhole increases hyperbolically with decreasing radius. Neglecting several actual influences on the casting procedure, it was found that for fibers with a diameter of 1 lm a pressure of 20 bar is necessary to overcome the negative capillary forces which hinder the melt entering a small hole owing to the bad wetting behavior of the melt on

13.2 Investment Casting

the mold. Considering real casting conditions, experiments showed that for parts with high aspect ratios, high pressure is beneficial even for parts with a diameter of 100 lm (see Section 13.5.6). However, it is very difficult to provide a pressure significantly higher than 4 bar because the pressure needs to be applied to the melt in a very short time interval t, where t ranges from 10–3 to 1 s. This means that a pressurized tank with a valve large in diameter is necessary to manage an immediate pressure increase. On the other hand, a high pressure is not always desired because with increasing pressure the surface roughness determined by the generally porous investment increases owing to better replication of the surface structure of the mold. Therefore, an ideally flat mold surface without pores would be required, which, however, prevents residual air from leaving the mold during the form filling process. 13.2.3 Centrifugal Casting

In centrifugal casting, the mold rotates and the melt fills the mold by centrifugal forces. Fig. 13-6 shows a scheme of the process. The mold is fixed horizontally or nearly horizontally at one end of the centrifugal arm. On the other side, a counterweight is fixed to keep the balance of the rotation system. In most centrifugal casting machines the melt is first poured into the runner system and then the system starts to rotate, but there also exist some machines where the melt is poured into the already rotating system. Here the centrifugal forces are higher because the machine rotates already at top speed and no further acceleration has to be taken into account. The centrifugal casting machine Ticast® (Kobelco Research Institute, Inc.) shown in Fig. 13-7 is based on this principle [46]. Here, in contrast to the scheme in Fig. 13-6, the casting metal is melted in a crucible fixed midway above the centrifugal arm. When the melting process is finished, the crucible is tilted and the melt is poured into a so-called drop mold located exactly in the middle of the rotating system. From here the melt flow is directed into the rotating mold and fills the mold owing to centrifugal forces. The rotational speed of centrifugal casting machines is in the range 350– 3000 rpm [48–50]. Modern centrifugal casting machines produce a higher pressure for form filling compared with vacuum pressure casting machines, which is beneficial for the casting of very small structures. However, the higher pres-

Fig. 13-6 Scheme of the centrifugal casting process

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Fig. 13-7 Centrifugal casting machine Ticast® by Kobelco Research Institute, Inc.

sure can give rise to various defects in the casting and the high turbulences may cause gas entrapment and favors gas porosity [51]. This disadvantages need to be weighed against the partly significantly improved form filling (see Section 13.5.6). 13.2.4 Example of a Typical Casting Process

After assembling the micropatterns on a wax sprue, the investment is mixed. For phosphate bonded investments a mixing ratio of 100 g of powder and 14– 23 ml of liquid is recommended [52]. With this mixture a total expansion of the mold between 1.2 and 2.4% can be achieved. It consists of the setting expansion due to chemical reactions in the investment and the thermal expansion during the heating (see Section 13.4.2). The expansion can be controlled additionally by the concentration of the liquid and can be varied in a wide range by adding distilled water to the liquid. Depending on the required mold size, powder amounts between 60 and 540 g are used in dental casting and in microcasting. The components are premixed manually for *30 s and then mixed for 60 s under vacuum. The processing time ranges between 4 and 6 min. During this time, the investment is poured into the cuvette containing the pattern assembly. The embedding process is supported by light vibrations using special equipment, for example Multivac® 4 by Degussa Dental GmbH. Subsequently the investment hardens for 20–50 min. According to the instructions for use in dental casting, the wax has to be eliminated in the next step by placing the mold in a preheated furnace at a temperature of about 300 8C. Depending on the mold size, this process takes between 30 and 60 min. Subsequently, the mold is transferred to another furnace for preheating with a starting temperature of 300 8C and is heated to the final temperature between 700 and about 1000 8C. Owing to the thermal expansion of the mold, a low heating rate of about 7 K/min is required. The final temperature needs to be held for about 20–60 min depending

13.3 Casting Alloys

on the mold size. Then the casting procedure begins with melting the alloy, for example in the vacuum pressure casting machine Prestomat® by Degussa Dental GmbH. Using Stabilor® G as casting alloy, the required amount of material (about 8–20 g) is melted at 1100 8C in a graphite crucible. After a 2 min dwell time, the preheated mold is placed in the machine and the casting process is started. Then the mold filled with the casting is removed, placed on a heat-resistant stone and is air-cooled. When the mold has reached ambient temperature, it is removed and the cast structure is cleaned in hydrofluoric acid and then washed in water. Finally, the microcomponents are separated from the runner and sprue.

13.3 Casting Alloys 13.3.1 Introduction

Although, in theory, all meltable metals can be cast, the casting suitability of alloys differs significantly. This is particularly important for microcasting, because here extremely small structures need to be filled sufficiently. Therefore, the most relevant factors for a casting alloy are high form filling ability and flowability of the melt. They are influenced by the viscosity of the melt, the wetting behavior of the form, the reaction with the mold and the atmosphere and, of course, by the solidification behavior. A high form filling ability and a good flowability are mainly guaranteed for precious alloys such as are used in jewelry and dental casting, for bronze (handcraft arts) and for high-strength alloys, such as CoCrMo alloys, which are especially used for dental castings. Steels are not widespread in microcasting because of their oxidation and corrosion sensitivity. 13.3.2 Gold Base Alloy as an Example of Precious Metals

In ancient times, precious alloys, especially gold and gold alloys, were used for casting small structures. The earliest example of a lost wax casting in gold alloy is a small animal figure on the rein-ring from Queen Pu-Abi’s chariot dated about 2600 BC [53, 54]. Further examples were recovered on the Greek islands, in Anatolia, in Mesopotamia and in the vicinity of the Dead Sea [54–56]. The gold coffin treasures of the Egyptians also include cast parts made by the lostwax method. The historical knowledge of forming gold by investment casting has now been rediscovered for the manufacturing of microparts. The advantage of gold and gold base alloys is their relatively low melting and casting temperatures. These materials are known to have excellent castability and a low oxidation affinity during the casting process. Because of their color and their high resistance against corrosion and tarnishing, gold and gold base

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alloys are eligible materials for jewelry. Good chemical properties and biocompatibility are essential for all materials used in oral health applications. Furthermore, sufficient mechanical properties regarding the service stresses are required [57]. For the fabrication of crowns and bridges, mainly alloys based on the ternary system AuAgCu are used. The increase in strength achievable compared with pure gold depends on the composition and the aging temperature. Scientific investigations of ternary AuAgCu alloys have shown a complex combination of constituent age-hardening mechanisms such as ordering processes of the CuAu I phase, precipitation and spinodal decomposition [58–60]. Similar results are documented for several commercial low-carat gold alloys for dental applications often containing Pd and some additional elements in minor amounts [61–63]. Owing to the good castability and high mechanical strength, an age-hardening gold base alloy was the first choice for research and development of microcasting based on investment casting. Especially the good casting properties recommended this material for replicating turbine parts with 10–30 lm wall thickness and specimens for mechanical testing with an aspect ratio (flow length-to-wall thickness ratio) up to about 30. Previous investigations led to the choice of the gold base dental alloy Stabilor® G, manufactured by Degussa Dental GmbH. Like the other age-hardening materials mentioned before, Stabilor® G is based on the ternary system AuAgCu and, therefore, shows good chemical resistance combined with excellent mechanical properties. It is especially suited for devices in mechanical engineering. Detailed specifications of this alloy concerning the chemical composition and the mechanical properties given by the supplier Degussa Dental GmbH are presented in Tables 13-1 and 13-2. The melting range is between 860 and 940 8C.

Table 13-1 Composition of Stabilor® G in wt.% (values given by manufacturer)

Au

Ag

Cu

Pd

Zn

Pt

Ir

58.0

23.3

12.0

5.5

1.0

0.1

0.1

Table 13-2 Mechanical properties of Stabilor® G (values given by manufacturer)

Condition

Yield stress (MPa)

Ultimate tensile strength (MPa)

Hardness HV5

Elongation to fracture (%)

As quenched Precipitation hardened

400 830

510 890

170 275

33 6

13.3 Casting Alloys

13.3.3 Bronze as an Example of Typical Casting Alloys

Other typical casting alloys such as bronzes, with 5–12 wt.% aluminum as primary alloying element, are especially used for corrosion-resistant castings because they form a thin oxide surface layer which protects the alloy against corrosion. In addition to aluminum, these alloys often contain iron, nickel or manganese in order to improve their mechanical properties. Commercial alloys based on such compositions rank among the best acid-resistant high-strength alloys [64]. Al-bronzes are excellent wear-resistant materials and are, therefore, used in gears, bearings, bushes etc. [65]. The microstructure of Al-bronzes depends mainly on the alloying elements and the cooling rate. Additionally, further microstructural changes occur during subsequent heat treatment. Benkisser et al. [66], for example, explained in detail the microstructure of a cast CuAl10FeNi alloy that underwent slow cooling. Faster cooling, however, results in non-equilibrium phases. In this context, Horn-Samodelkin et al. reported the presence of a martensitic phase in a sand cast plate of G-CuAl10Ni [67]. For the binary system Cu–Al, Fig. 13-8 shows a transition diagram [68] with both the equilibrium phases and the transition temperatures for order and martensitic transformations. After solution heat treatment in the b-region and subsequent quenching in water, the b-phase still exists at ambient temperature. During quenching, the alloy passes two through order transformations and later forms martensite. The modification of the martensite phase depends on the Al content of the alloy. Additionally, three different martensites, a', b 01 and c', can be distinguished by their crystal structures. The advantage of Al-bronze for microcasting is the formation of a thin, transparent oxide layer, which already gives the casting a metal brightness without any chemical cleaning or mechanical polishing. Furthermore, compared with Sn-bronze, Al-bronze has a narrow solidification interval between 1020 and 1040 8C, which limits the tendency to segregate during solidification. Like Al-

Fig. 13-8 Transition diagram for binary CuAl alloys. Continuous line, phase diagram; dashed line, transition temperatures of order transformations and martensite formation. Adapted from Kainuma et al. [68]

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13 Microcasting Table 13-3 Composition of Al-bronze in wt.% (according to DIN 1714)

Cu

Al

Ni

Fe

76

8.5–11.0

4.0–6.5

3.5–5.5

Table 13-4 Mechanical properties of Al-bronze (according to DIN 1714)

Yield stress (MPa)

Ultimate tensile strength (MPa)

Hardness HV5

Elongation to fracture (%)

270–300

600–700

140–160

12–14

bronzes, the CuAl10Ni alloy used for our investigations shows generally good mechanical properties and good slip properties and is resistant against sea water. The chemical composition and the mechanical properties according to DIN 1714 are shown in Tables 13-3 and 13-4. 13.3.4 CoCrMo Alloy as an Example of High Strength Materials

CoCrMo alloys are well-known alloys in medical and dental applications. Especially in surgical implant applications, Co-base alloys are used as materials for the reconstruction of artificial joints because of their good mechanical properties and their excellent corrosion and biodegradation resistance. In dental techniques, CoCrMo alloys are mainly used for casting dentures. The high ultimate tensile and fatigue strength of the alloy combined with a sufficient elongation at fracture are of great importance [69]. CoCrMo alloys also guarantee biocompatibility, which means that the material itself remains unchanged by body liquids and does not interact unfavorably with body tissue [70]. The superior corrosion resistance is due to the fact that alloys with high chromium content form a passive oxide layer on the surface, protecting the alloy against the environment of the body. Cast CoCrMo alloys are known to exhibit an inhomogeneous, large-grained microstructure [71]. SEM and TEM investigations showed that the apparent two-phase microstructure seen in optical microscopy consists of many more phases. Investigations of a cobalt base alloy with 0.4% C, 30% Cr and 5% Mo, using energy-filtered TEM, proved the existence of an extremely fine ternary eutectic additionally to a blocky r-phase [72]. According to Kulmburg et al., it consists of a cobalt-rich solid solution, M23C6- and M6C-carbides. The extremely fine ternary eutectic is thought to be the reason for the good casting properties of the material. Other groups investigating a Co base alloy with 27.0–30.0% Cr, 5.0–7.0% Mo and maximum 1.0% Ni, 0.75% Fe, 0.35% C, 1.0% Si and 1.0% Mn found cobalt-rich dendrites and an interdentritic four-phase mixture containing a cobalt-rich c-phase, a chromium-rich M23C6 phase (with M = Co, Cr or Mo), an M7C6 phase and a chromium and molybdenum-rich r-phase [73, 74].

13.4 Investment Materials Table 13-5 Mechanical properties of Biosil® f (values given by manufacturer)

Yield stress (MPa)

Ultimate tensile strength (MPa)

Hardness HV10

Elongation to fracture (%)

Young’s modulus (GPa)

700

900

400

5

220

Table 13-6 Composition of Biosil® f in wt.% (values given by manufacturer)

Co

Cr

Mo

Si

Mn

C

64.8

28.5

5.3

0.5

0.5

0.4

Often these interdendritic phases reduce the ductility and the corrosion resistance and also limit the strength. Therefore, a subsequent solution heat treatment is recommended for cast CoCrMo alloys. According to Clemow and Daniell [73], the optimum temperature for this is 1225 8C. Exact temperature control during heat treatment is of great importance because the annealing needs to be kept below the eutectic temperature of 1235 8C in the CoCrMo system. According to Sullivan et al. [75], for CoCrMo alloys a yield stress of at least 450 MPa, an ultimate tensile strength of 655 MPa and an elongation at fracture of 8% can be expected. Interdendritic porosity, often found in cast CoCrMo alloys, is detrimental to the mechanical properties. The dental CoCrMo alloy Biosil® f used for microcasting reaches a nominal yield stress of 700 MPa, an ultimate tensile strength of 900 MPa and an elongation at fracture of 5% according to the manufacturer (Table 13-5). The exact chemical composition of the alloy is given in Table 13-6. The melting temperature of the alloy lies between 1320 and 1380 8C. The recommended preheating temperature for this alloy is 1000 8C and the recommended casting temperature 1500 8C [76]. During casting, a passivation layer is formed on the surface which should not be destroyed. Therefore, the cast part should not be treated with a stripping agent [76].

13.4 Investment Materials 13.4.1 Introduction

The investment is a ceramic slurry in which the pattern is embedded to produce a mold for casting. It has a significant influence on the casting result and should therefore meet several requirements. For embedding purposes, the investment should have good flowability and processibility. After drying, further

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burning at increasing temperatures is necessary in order to guarantee the required mechanical and thermal stability. The resulting mold should be chemically inert with regard to the metal melt, and furthermore it should show an adequate dimensional behavior, sufficient mechanical strength and low surface roughness and porosity [77]. At the end of the casting process, easy removal of the mold from the cast part is desirable. The investment is a mixture of a binder and a filler, with the binder itself consisting of powder and liquid. The solidification of the investment is facilitated by a chemical reaction of the binder components which results in stable inorganic phases. Most common binder materials used in dental and jewelry casting are slurries or liquids based on phosphate, plaster or silicate. For the shell casting technique, which is mainly applied for macroparts, often binders on a silicate base are used. Fillers are in all cases ceramic powders. Typical minerals used as fillers in investments are quartz, cristobalite, aluminum oxide, zirconium oxide, zirconium silicate and burned potter’s clay minerals such as mullite and molochite. In dental techniques, the filler materials mainly used are quartz and cristobalite, which are both modifications of SiO2. Additionally, for investment casting, the amorphous modification of SiO2 named melt quartz or amorphous quartz is used. The special importance of the minerals quartz and cristobalite lies in their high thermal expansion behavior. In contrast to amorphous quartz, these minerals show phase transitions which are associated with a strong volume increase. The transition from low- to high-temperature quartz at 573 8C results in a volume increase of 0.8%. Cristobalite, which at ambient temperature has a 14 times higher specific volume than quartz, shows a volume increase of 2.8% on transforming from the low- to high-temperature form at a transition temperature of 270 8C [78]. Owing to kinetic effects, impurities and mismatches in the crystal structure of the raw material, the real values of the volume increase may differ from the values given in the literature. Fig. 13-9 compares the linear thermal expansion of the three SiO2 modifications quartz, cristobalite and amorphous quartz [79]. Here, cristobalite shows a sudden change from the low- to the high-temperature modification in the vicinity of the transition temperature (*270 8C). The total linear thermal expansion of cristobalite is 1.6% for temperatures up to 800 8C. However, quartz expands by 1.3% up to temperatures of about 650 8C, whereas at higher temperatures a contraction is observed. The thermal expansion of amorphous quartz is very low, < 0.1%. Hence the expansion behavior of the investment during heating to standard mold temperatures between about 700 and 1000 8C is mainly controlled by its cristobalite and quartz content. With increasing cristobalite content, the expansion of the investment increases. For high-precision replication, good contour accuracy is required in addition. This is guaranteed by special fine-scaled investments. However, excessive downscaling of the filler particle size is problematic, because the increasing particle surface requires a higher liquid content. This results especially for phosphate bonded investments in drying cracks during hardening. For plaster bonded in-

13.4 Investment Materials

Fig. 13-9 Linear thermal expansion of quartz, cristobalite and amorphous quartz. Adapted from Degussa Dental GmbH [79]

vestments, however, the amount of water can be varied over a wider range without inducing cracks, but it has to be stressed that an increasing water content in the investment hampers the drying process [17]. 13.4.2 Phosphate Bonded Investments

The choice of the investment material depends on the cast metal and the required strength of the mold. Phosphate bonded investments were initially used for dental alloys with a high melting temperature at casting temperatures between 1200 and 1500 8C. However, in dental casting, these investments are nowadays used for all alloys because of their high heat resistance, good mechanical strength and convenient workability. They consist of magnesium oxide and ammonium hydrogenphosphate as binder and the two different SiO2 modifications quartz and cristobalite as filler. Special investments for gold base alloys sometimes also contain graphite powder in order to produce a reducing atmosphere in the mold. The powdery binder and filler components are mixed with a liquid which mainly consists of aqueous silica sol. The water content in the liquid is necessary to facilitate a chemical reaction [80]: NH4 H2 PO4 ‡ MgO ‡ 5H2 O

!

MgNH4 PO4
6H2 O

In the burning process, ammonium phosphate is converted into magnesium pyrophosphate, releasing water and ammonia according to [80] MgNH4 PO4
6H2 O 2…MgNH4 PO4
H2 O†

! !

MgNH4 PO4
H2 O ‡ 5H2 O Mg2 P2 O7 ‡ 3H2 O ‡ 2NH3

> 1608C > 2508C

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13 Microcasting

Above 1000 8C the investment decomposes by continuous reaction of excrescent magnesium oxide with the phosphate of the binder and the silicate of the filler [80]: Mg2 P2 O7 ‡ MgO 2MgO ‡ SiO2

!

!

Mg3 …PO4 †2

Mg2 SiO4

> 10008C > 10008C

Commercial investments for dental casting are designed in such a way that they exactly compensate the shrinkage of a given metal during solidification. During the manufacturing process, the necessary expansion of the investment is mainly achieved by the expansion of the filler due to the change from a lowto a high-temperature phase (cristobalite, 270 8C; quartz, 573 8C; see Fig. 13-10), which comprises a modification of the crystal structure. To allow enough time for the phase transitions, a slow heating regime is required in order to avoid cracks. Degussa Dental GmbH recommends a heating rate of 7 K/min for several of their phosphate bonded investments [52]. The thermal and chemical expansion of the binder is controlled by the silica sol concentration of the liquid and generally increases with increasing concentration of the liquid (Fig. 13-10). Additionally, the amount of silica sol determines the surface roughness of the mold because it transforms into amorphous SiO2 during burning and fills the pores of the mold. Whereas silica sol reduces the porosity of the mold, adding graphite increases it. Graphite-containing investments are used for better gas permeablity of the mold and easier deflasking of the cast parts. Phosphate bonded investments have a higher strength than plaster bonded investments [81] and are more convenient to handle. Hence they are widespread in dental laboratories for precious metals and for model casting alloys on a

Fig. 13-10 Thermal expansion behavior of phosphate bonded investment. Adapted from Degussa Dental GmbH [79]

13.4 Investment Materials

CoCr base. For the latter, additionally the higher possible preheating temperatures up to 1000 8C are important. 13.4.3 Plaster Bonded Investments

When casting microparts, a great difficulty is the removal of the investment. For gold base alloys, phosphate bonded investment can be removed by acid owing to the higher chemical resistance of the metal compared with the investment, but for base alloys such as bronze or CoCrMo alloys this procedure is not applicable. Hence special investments which are easily soluble are required. This is achieved by using plaster as binder. Like phosphate bonded investments, plaster bonded investments contain quartz and cristobalite as refractory filler materials. Additionally, the investments often contain auxiliary materials. The material is commonly employed in casting gold alloys with high gold content and with liquidus temperatures not higher than 1080 8C [81]. In dental applications, plaster bonded investments are not used as often as in jewelry casting owing to the danger of sulfur release by reactions of the melt with the investment and hence resulting low strength of the casting. In jewelry and artistic casting, however, most investments are plaster based, because they achieve a low surface roughness and can be easily removed also from complicated-shaped structures. Plaster used in investments is hard plaster or super-hard plaster, chemically both a-semihydrates. For cuvette embeddings, hard plaster can be used, but if precision and high mechanical strength are required then super-hard plaster should be adopted [82]. When mixing the slurry for embedding, the chemical reaction of a-semihydrate with water results in calcium sulfate dihydrate [83]: 1 3 CaSO4
H2 O ‡ H2 O 2 2

!

CaSO4
2H2 O

This rehydration is an exothermic process. Stoichiometrically, 18.7 ml of water are needed to rehydrate 100 g of plaster, but for mixing a well-flowing slurry for embedding additional water is required [82]. Depending on the particle size and form of the crystals, 100 g of a-semihydrates need 28–32 or 19–29 ml of water, respectively, depending on the modification hard plaster or super-hard plaster. During burning, the crystal water is released and the anhydrite of calcium sulfate for the mold is obtained [79]: CaSO4
2H2 O

!

CaSO4 ‡ 2H2 O

Above 750 8C, calcium sulfate decomposes according to [79] 2CaSO4

!

2CaO ‡ 2SO2 ‡ O2

The generation of sulfur dioxide may result in deterioration of the casting due to gas bubbles in the near-surface region if the gas cannot be dissipated

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through the porous mold. Another problem with plaster bonded investments is the decomposition of the investment between 500 and 600 8C if carbon is present [79]: CaSO4 ‡ 4C

!

CaS ‡ 2CO2

During setting, the plaster shows a change in volume. Although the real volume will decrease by 7% when the starting substances are mixed together, the apparent volume will increase owing to the feeding by the slurry during the solidification of the mold. The volume change is determined by the linear difference in elongation. For dental plasters, first a linear contraction is found followed by a linear setting expansion between 0.05 and 0.15% for special hard plasters [82]. 13.4.4 Influence of the Investment on the Surface Roughness

The investment has a significant influence on the surface roughness of the cast part. With increasing form filling ability of the casting alloy, the surface roughness of the cast part approaches more and more that of the surface of the mold. Therefore, a very smooth surface of the mold with only a few pores is necessary. Three different ways are possible to meet these requirements: · coating the pattern with an extremely fine ceramic; · infiltrating the sintered mold with a ceramic suspension; and · modifying the investment by addition of fine ceramic particles [84]. Within the scope of the Collaborative Research Center 499 [85], comparative roughness measurements were made employing the three methods mentioned above. The surface topography of the cast parts, which replaced or replicated a smooth, injection-molded PMMA plate 4 ´ 6 mm2 in size, was imaged by a tactile system (Perthometer®, Mahr GmbH) with a tip radius of the needle of about 1.4 lm and a load between 10–3 and 10–5 N. The roughness values Ra and Rmax were calculated along a 5.6 mm long segment on the plain side of each sample. Ra is the arithmetic mean of the deviations of the absolute values with respect to the center line of the roughness profile [86–88]. Rmax is defined as the largest single deviation from peak to valley along the measured segment. The surface roughness of the ceramic mold was determined by a contactless, optical-measuring system (Microglider®, Fries Research & Technology GmbH) because the mold is too weak to withstand a tactile system. In the following, the three different methods for reducing the surface roughness of the ceramic mold and therefore of the cast part will be described.

13.4 Investment Materials

a)

b)

c)

d)

e)

f)

Fig. 13-11 Sketch of casting process with coated pattern. a) PMMA pattern; b) coated pattern; c) coated pattern in ceramic slip; d) hollow form after sintering with coating left inside the mold; e) metal-filled form; f) cast part

13.4.4.1 Coating the Pattern Investigations on surfaces improved by coating the pattern were carried out with a 1% aqueous methylcellulose solution containing 5% fine-scale SiO2 powder. The pattern was dipped into the liquid and dried on air. Subsequently, the coated pattern was embedded in an investment slip (Deguvest® CF by Degussa Dental GmbH), dried and burned as described before. The hollow form was then filled with the molten gold base alloy Stabilor® G in a vacuum pressure cast machine (Prestomat® by Degussa Dental GmbH). The applied pressure was 4 bar. A sketch of the process is given in Fig. 13-11. The profile for surface roughness for cast parts that were replicated in the commercial investment shows large amplitudes with roughness values of Ra = 1.13 lm and Rmax = 8.41 lm whereas a coated specimen achieves values of only Ra = 0.74 lm and Rmax = 6.19 lm. Therefore, coating of the mold is beneficial to a low surface roughness of the cast part. However, it must be stressed that coating a plastic or wax pattern is difficult owing to the poor adhesion of the liquid on the pattern. Dipping the PMMA pattern in the slurry once leads to a coating which delivers relatively good results. Repeated dipping two or three times, however, often results in flaking of the coating.

13.4.4.2 Infiltrating the Mold Infiltration was examined using suspensions of fine-grained SiO2 powder stired in distilled water (1 : 10). In further experiments it was diluted to 1 : 20 and 1 : 40. The process of infiltration is illustrated in Fig. 13-12. First the mold was manufactured as discussed earlier. The hollow form was then infiltrated with the suspension and sintered in a second burning process followed by the final casting process. Compared with the coating technique, the infiltration of a microstructured sintered form is much more difficult. If the viscosity of the suspension is too high, it does not infiltrate the pores of the mold but instead forms a thick layer

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a)

b)

c)

d)

e)

f)

Fig. 13-12 Sketch of casting process with infiltration. a) PMMA pattern; b) embedded in ceramic slip; c) hollow form; d) infiltrated investment after second sintering; e) metal-filled form; f) cast part

on the surface, leading to rounded edges and differences in tolerance. However, when the viscosity of the infiltration liquid is too low, then there is no visible effect on the mold. It fully penetrates the porous investment without leaving a layer on the mold surface. Additional investigations showed that it is almost impossible to realize a homogeneous infiltration in a closed mold with internal microstructures. Appropriate tests for this technique proved inconclusive.

13.4.4.3 Modifying the Investment Reduction of the surface roughness is also possible by modifying the investment compound itself. Therefore, different amounts of fine-grained SiO2 powder were added to the Deguvest investment. To maintain the consistency of the slip, the liquid content was increased at the same time. The modifications used are summarized in Table 13-7. When the commercial investment (Deguvest® CF) was modified, much better results than with the first two methods were achieved. Compared with the commercial investment, modification 1 results in a slightly lower Ra value, but in an increased Rmax value for the cast parts (Table 13-8). The second modification, however, achieves a significant reduction in both surface roughness values (Table 13-8). The Ra value decreases from 1.13 lm for the commercial investment to 0.44 lm for modification 2 and at the same time the Rmax value decreases from 8.41 to 3.24 lm. Table 13-7 Composition of the investment modifications used

Deguvest® CF SiO2 powder Liquid

Modification 1

Modification 2

Modification 3

60 g 12 g 10 ml water + 30 ml liquid

60 g 12 g 40 ml

60 g 8g 29 ml

13.5 Cast Microparts and Their Properties Table 13-8 Influence of investment on the roughness of cast surfaces

Investment

Ra (lm)

Rmax (lm)

Commercial investment Modification 1 Modification 2

1.13 0.97 0.44

8.41 10.74 3.24

The lower surface roughness of the parts cast in modified investment originates in a reduction of the pore size in the mold [84] owing to the use of a larger amount of fine-grained filler in the investment. However, there will be a limitation to this development owing to the necessity for an open porosity in the mold in order to allow the exit of gas out of the cavities during the form filling process. Another problem when using fine-grained investment is the increasing demand for liquid for mixing the slurry owing to the large surface area of the fine-grained powder that needs to be wetted. The high liquid content in the investment often causes cracks in the mold during drying and burning.

13.5 Cast Microparts and Their Properties 13.5.1 Examples of Cast Microparts

Within the scope of this research, a wide variety of microparts were cast, including specimens for mechanical testing (bending bars, tensile test specimens) and parts for a demonstrator consisting of a turbine and a gear unit. Fig. 13-13 (left) displays an array of bending bars with a geometry of 200 ´ 200 ´ 1800 lm3 each.

Fig. 13-13 Specimens for mechanical testing cast in gold base alloy Stabilor® G. Left: bending bars (200 ´ 200 ´ 1800 lm3); right, tensile test specimens (260 ´ 130 ´ 4000 lm3) (top: plastic pattern; bottom: cast part)

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Fig. 13-14 Turbine parts cast in gold base alloy. a) Housing; b) nozzle plate; c) (blade) wheel

Fig. 13-15 Cast turbine parts made of different alloys. a) Al-bronze; b) CoCrMo alloy

On the right, a tensile test specimen with a total length of 4 mm and a rectangular cross-section of 130 ´ 260 lm2 in the gage length is shown below the plastic pattern used as a model. The discussed specimens for mechanical testing are both made of Stabilor® G. In addition, microparts for a demonstrator consisting of a turbine with a diameter of 4 mm and a gear unit were cast in the gold base alloy Stabilor® G. The smallest geometry can be found on the nozzle plate (Fig. 13-14 b) with a channel width of 25 lm and a channel height of 200 lm. The blades of the wheel (Fig. 13-14 c) have a thickness of 100 lm, a length of about 400 lm and a structural height of 190 lm. The housing (Fig. 13-14 a) has holes with a diameter of 200 lm in the outer circle and 210 lm in the inner circle. The entire height of the housing is about 395 lm and 245 lm in the inner area where the nozzles are located. The dimensional accuracy is discussed in Section 13.5.3. Additionally to the gold base alloy, a bronze and a CoCrMo alloy have been used to cast microturbine parts. Fig. 13-15 shows corresponding nozzle plates with an outer diameter of 1.9 mm. The channel width is 75 lm and the height is 350 lm. The nozzles themselves have a diameter of 200 lm.

13.5 Cast Microparts and Their Properties

13.5.2 Microstructure/Grain Size

The microstructure of cast parts is mainly influenced by the mold temperature. An increase in the mold temperature directly lowers the cooling rate. For a quantitative investigation, wax sticks with a diameter of 3 mm were equipped with three thermocouples each: one near the sprue bottom (TC 1), one in the middle (TC 2) and one at the top of the stick (TC 3). A single wax stick carrying the thermocouples was then embedded in an investment. After heating to 700 8C, the mold was kept constantly at this temperature. This process was repeated with a second wax stick–investment combination that was allowed to cool after burning and then kept at 100 8C for 1 h. Later, both molds were filled with the gold base alloy Stabilor® G using a casting temperature of 1100 8C. Fig. 13-16 shows the time-dependent temperature development in the mold during casting. The cooling rate is highest at the beginning of the casting process. The temperature for a mold that was kept at constant 700 8C (Fig. 13-16 b) decreases by only 150 8C within the first 20 s, whereas for a mold held at 100 8C (Fig. 13-16 a) a much higher cooling rate is observed. Grain size analyses of the above sticks show a significant influence of the mold temperature on the average grain size (Table 13-9). The grain size increases from 24 lm for a mold tem-

a)

Fig. 13-16 Influence of two mold temperatures on the temperature development in a cast stick made of gold base alloy Stabilor® G with a diameter of 3 mm. (a) Mold temperature 100 8C; (b) mold temperature 700 8C. TC 1, thermocouple at the gate; TC 2, thermocouple in the middle of the stick; TC 3, thermocouple at the top end

b)

383

384

13 Microcasting Table 13-9 Influence of mold temperature on grain size for sticks of 3 mm diameter cast in the gold base alloy Stabilor® G

Mold temperature (88 C)

Grain size (lm)

100 400 700 1000

24 35 56 91

Table 13-10 Influence of mold temperature on grain size for sticks of 3 mm diameter cast in Al-bronze

Mold temperature (88 C)

Grain size (lm)

100 400 700 1000

*30 *45 *100 *170

Fig. 13-17 Influence of mold temperature on the microstructure for sticks with 3 mm diameter cast in Al-bronze. a) Mold temperature 100 8C; b) mold temperature 10008C

perature of 100 8C up to 91 lm at 1000 8C. However, while a stick diameter of 3 mm is still a macrosize, smaller grain sizes have been found for microstructures (see examples in Chapter 19). They generally underlie higher cooling rates as a result of the smaller amount of heat induced in the mold during casting. The influence of the mold temperature will increase with decreasing structure size because the influence of the heat content of the mold increases relative to the heat content of the cast part [37]. In addition to the grain size, the cooling rate also influences the precipitation behavior of hardened alloys such as Stabilor® G because with an increasing cooling rate the solidification turns more and more into a quenching process far away from any metallurgical equilib-

13.5 Cast Microparts and Their Properties

rium. Therefore, when using low mold temperatures a subsequent annealing process is recommended in order to achieve an adequate age-hardened microstructure. The effect of the mold temperature or cooling rate on the grain size of Albronze is shown in Table 13-10. This material, in comparison with the gold alloy (Table 13-9), exhibits distinctly coarser grain sizes. For cast sticks with a diameter of 3 mm, the grain size increases from about 30 lm for a mold temperature of 100 8C up to about 170 lm at 1000 8C. Additionally, cross-sections of the Al-bronze sticks in Fig. 13-17 show an obvious coarsening of the different phases with increasing mold temperature from 100 to 1000 8C. 13.5.3 Dimensional Accuracy

In the microcasting process, very good dimensional accuracy can be achieved when using an investment which is in compliance with a casting metal regarding the expansion behavior. In order to determine the dimensional accuracy of the casting process, an array of 144 vertical bending bars was embedded in Deguvest® CF, the investment for gold base alloys supplied by Degussa Dental GmbH and replicated with Stabilor® G as corresponding metal. In contrast to the standard microcasting process, a wax pattern replicated (via silicon) from a milled brass array was used. Fig. 13-18 shows the mean width of the bending bars for the original made of brass, for the wax pattern and for the cast parts averaged over about 100 specimens each. The original, with 209.2 lm, is distinctly larger than the nominal value (200.0 lm). The size of the replications decreases slightly from the wax pattern (207.8 lm) to the cast bending bars (206.3 lm). For the precision of the casting process, however, only the dimensional accuracy of the cast parts compared with the wax pattern is of interest. As the standard deviations are similar at 4.2 and 3.9 lm, respectively, it is therefore possible to compare the values for the mean widths. In the given example, the difference is only 1.5 lm, which is within the accuracy of measurement. This value displays a very good dimensional accuracy of the casting process.

Fig. 13-18 Dimensional accuracy of cast parts made of Stabilor® G bending bars 200 ´ 200 ´ 1800 lm3 cast as array of 144 specimens at a mold temperature of 700 8C. Shown is the mean width and the standard deviation r of the original, the wax patterns and the cast part; all values given in lm

385

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13 Microcasting

13.5.4 Surface Roughness

The surface roughness is influenced by various parameters. One parameter is the investment, as discussed in Section 13.4.4; the two other important parameters are the mold temperature and the casting pressure. The influence of these parameters was measured on tensile test specimens made of the gold base alloy Stabilor® G using both a confocal white light microscope (NanoFocus lSurf NanoFocus AG) and a scanning optical system (Microglider®, Fries Research & Technology GmbH). Both set-ups are contactless measuring systems. The measuring lengths are comparable and lie between 1.2 and 1.6 mm. Tables 13-11 and 13-12 present the roughness values Rt and Rz measured by NanoFocus and Ra and Rz measured by FRT for specimens cast under different conditions. Comparing specimens cast at a mold temperature of 1000 8C but with different casting techniques, it can be seen that the surface roughness achieved by vacuum pressure casting is significantly lower than that by centrifugal casting. This is due to the different filling pressures applied in the two methods. Whereas the centrifugal casting machine works with a filling pressure of about 20– 25 bar, the vacuum pressure casting machine fills the mold with only 4 bar. This results in a different quality of replication of the surface roughness of the mold. The high pressure in the centrifugal casting machine leads to a better replication of the mold surface but at the same time results in a higher surface rough-

Table 13-11 Rt and Rz values measured with NanoFocus [89]. CC: centrifugal casting; VPC: vacuum pressure casting; measuring length: *1.6 mm

Mold Casting temperature technique (88C)

No. of specimens

700 1000 1000

63 36 11

CC CC VPC

Rt (lm)

Rz (lm)

Mean value Standard deviation

Mean value Standard deviation

9.53 8.68 4.19

5.96 5.34 2.85

4.51 3.82 0.50

3.13 2.63 0.27

Table 13-12 Ra and Rz values measured with FRT. CC: centrifugal casting; VPC: vacuum pressure casting; measuring length: 1.2–1.5 mm

Mold Casting temperature technique (88C)

No. of specimens

700 1000 1000

6 12 5

CC CC VPC

Rt (lm)

Rz (lm)

Mean value Standard deviation

Mean value Standard deviation

1.35 1.12 0.78

10.27 8.03 4.93

0.48 0.56 0.14

2.32 4.15 0.69

13.5 Cast Microparts and Their Properties

387

ness or worse surface quality, respectively. In comparison, the lower filling pressure of the vacuum die cast machine produces significantly lower surface roughnesses. On the other hand, this also causes problems during form filling. Hence the investigator needs to find the optimum for each microstructure balancing the highest necessary filling pressure in order to fill the microcavities against the lowest possible pressure in order to achieve the least surface roughness. Furthermore, Tables 13-11 and 13-12 show that a higher mold temperature of 1000 8C also has a positive influence on the surface roughness. It is thought that the formation of a thin oxide layer during solidification of the melt causes a smoother surface of the specimens. Therefore, a mold temperature of 1000 8C is beneficial for the form filling and also for a low surface roughness. 13.5.5 Mechanical Properties

In order to understand the relationship between process, microstructure and mechanical properties, tensile tests with microscaled specimens made of a gold base alloy were carried out at the Institute for Materials Science I, University of Karlsruhe, Germany. The specimens have a thickness of 130 lm and a rectangular cross-section with a width of 260 lm. A detailed overview of all results is given in Chapter 19. Here, only the most important results with respect to the casting parameters are presented. This includes the yield stress and the ultimate tensile strength as a function of the mold temperature, the aging condition and the casting procedure (Table 13-13). For centrifugal casting, an increase in the yield stress and the ultimate tensile stress with increasing mold temperature is observed. A mold temperature of 400 8C, for example, results in a yield stress of 396 MPa. The yield stress increases to 518 and 521 MPa for higher mold temperatures of 700 and 1000 8C, respectively. At the same time, the ultimate tensile stress reaches 628 MPa for a mold temperature of 400 8C and 824 and 776 MPa for the higher mold temperatures. Therefore, in order to achieve a high mechanical strength, for a gold base alloy, a mold temperature between 700 and 1000 8C is required. It has to be

Table 13-13 Mechanical properties of cast Stabilor® G as a function of mold temperature, aging condition and casting procedure a

Yield stress (MPa) Ultimate tensile strength (MPa) No. of specimens

Tmold = 400 8 C; CC

Tmold = 700 8 C; CC

Tmold = 1000 8 C; CC

Tmold = 1000 8 C *; CC

Tmold = 1000 8 C; VPC

396 628

518 824

521 776

716 1005

354 594

9

8

10

3

9

a CC: centrifugal casting; VPC: vacuum pressure casting; * heat treated (5 h, 800 8C, vacuum).

388

13 Microcasting

stressed, however, that these values of mechanical properties are distinctly lower than those given by the manufacturer (measured on macroparts; compare Section 13.3.2). Besides a possible influence of the grain size, this is mostly due to the aging behavior of the cast parts during the cooling phase, because the gold base alloy is a precipitation hardening material which needs to cool slowly. However, for microparts the amount of heat induced by the melt during the casting process is not high enough to generate significant aging while cooling. Therefore, an additional aging process is necessary to achieve full hardening. This can be seen in the fourth column of Table 13-13. The specimens cast in a mold at 1000 8C show an increase in mechanical strength after being heat treated for 5 h at 800 8C in an evacuated quartz tube with subsequent slow cooling inside this set-up. However, the yield strength of the aged microspecimens is distinctly lower than that of their macro specimens while the ultimate tensile stress significantly exceeds the values for macroparts significantly. Another interesting result is derived from the comparison of the two casting techniques – centrifugal casting and vacuum pressure casting – with respect to the mechanical properties achieved. The centrifugal cast specimens show a yield stress of 521 MPa and an ultimate tensile strength of 776 MPa for a mold temperature of 1000 8C, whereas the vacuum pressure cast specimens show a yield stress of 354 MPa and an ultimate tensile strength of 594 MPa. Here, the much larger grain size of the vacuum cast specimens seems to dominate the results. Investigations reported in Chapter 19 indicate a grain size of 64 lm for these specimens compared with about 10 lm for the centrifugal cast specimens. 13.5.6 Achievable Structure Size, Flow Length and Aspect Ratio

To a great extent, the smallest achievable structure size depends on the aspect ratio, which is defined as the ratio of flow length to wall thickness. Investigations showed that very small structures can be cast with the gold base alloy owing to its good flowability and form filling behavior. Wall structures down to 20 lm width were produced with an aspect ratio of 6. If adequate patterns were available, even smaller structures should be castable with this alloy. The smallest structures cast with all three alloys in this project, the gold base alloy Stabilor® G, the Al-bronze and the CoCrMo alloy Biosil® f, are channels on a turbine plate with a width of 25 lm and a structure height of 350 lm. The flow length and aspect ratio achievable are mainly influenced by the preheating temperature of the ceramic mold and by the filling pressure. Flow length tests were carried out in order to determine the dependence of these two parameters on the form filling behavior. Therefore, two different fiber patterns were produced which consist of 10 organic fibers each mounted on a horizontal wax sprue. In the first case the fibers had a diameter of 210–230 lm and in the other case the diameter was 100 lm. The fibers were cast in the gold base alloy Stabilor® G and in Al-bronze using different mold temperatures and filling pressures. The results of the flow length tests are presented in Fig. 13-19. For

13.5 Cast Microparts and Their Properties

a)

c)

b)

d)

Fig. 13-19 Flow lengths of different metal melts during form filling of Stabilor® G in holes with different diameters in the submillimeter range. a) Stabilor® G (230 lm); b) Stabilor® G (100 lm), c) Al-bronze (210 lm), d) Al-bronze (100 lm)

Stabilor® G, with a fiber diameter of 230 lm (Fig. 13-19 a) and a mold temperature of 1000 8C, which exceeds the liquidus temperature of 940 8C, the maximum flow length of about 9 mm is reached at a relatively low pressure of 2 bar. For a mold temperature of 700 8C, a linear increase in flow length with increasing pressure is observed. Here, a pressure of 4 bar – the value where the vacuum pressure casting machine is at its limit – results in a flow length of nearly 6 mm. For a fiber diameter of 100 lm (Fig. 13-19 b), the same value can be reached at a pressure of 25 bar using the centrifugal casting machine. In summary, for a mold temperature of 700 8C an increase in pressure from 4 to 25 bar allows the fiber diameter to be halved while achieving the same flow length. This corresponds to an aspect ratio of 60. For the Al-bronze, there is a more pronounced difference in achievable flow lengths dependending on the fiber geometry. Whereas for fibers with a diameter of 210 lm (Fig. 13-19 c) the flow length increases similarly for the two mold temperatures with increasing pressure, for fibers with a diameter of 100 lm

389

390

13 Microcasting

(Fig. 13-19 d) a flow length of only about 2 mm was reached, resulting in an aspect ratio of 20. Hence the improvement regarding the flow length achievable by increasing the pressure is very small. The reason for this behavior is the distinctly higher solidification temperature of Al-bronze compared with the gold base alloy Stabilor® G. Consequently, the much larger difference between mold and solidification temperature for the Al-bronze yields a much faster solidification of the bronze melt when entering the mold compared with the melt of the gold base alloy and, therefore, an increase in pressure has only a small influence on the achievable flow length of the Al-bronze. In addition, the different solidifications of the two alloys, Al-bronze and Stabilor® G, should be taken into account. Whereas the gold base alloy is a solid solution with very fine precipitations, the Al-bronze forms a much coarser multiphased microstructure which is likely to be detrimental to the form filling of extremely fine structures. It is therefore suggested that, compared with Stabilor® G, the use of Al-bronze should be limited to microstructures with lower aspect ratios than for Stabilor® G. However, these limitations might only cause problems for extremely fine structures with aspect ratios of > 20 and diameters < 100 lm. Current research has shown that all microstructures mentioned in this chapter can be cast with Al-bronze. Even the tensile test specimens with an aspect ratio of 30 and a wall thickness of only 130 lm were castable with this material, because their cross-sectional area (33 800 lm2) is comparable to that of fibers with a diameter of 210 lm (34 636 lm2). Therefore, microcasting is an ideal processing tool for the replication of nearly all microstructures with large flow lengths or large aspect ratios, even if the wall thicknesses reach a minimum value.

13.6 Conclusions

Investment casting is a suitable technique for the manufacture of metallic microparts. The smallest channel structures cast with all three alloys tested, a gold base alloy (Stabilor® G), an Al-bronze alloy and a CoCrMo alloy, have a width of 25 lm and a height of 350 lm. In addition, with Stabilor® G even smaller wall structures 20 lm wide and 120 lm high could be cast. Nevertheless, the dimensional limits for microcasting have not been reached yet. Especially gold base alloys such as Stabilor® G offer the potential for further miniaturization of investment cast components. The results of microcasting depend on several parameters such as filling pressure and mold temperature which are of special importance. It has been shown that with our equipment centrifugal casting provides a significantly higher filling pressure than vacuum pressure casting. Thus centrifugal casting is of advantage when long and thin fiber-like parts are to be cast. The mold temperature has a significant influence on the average grain size of cast specimens. The grain size of vacuum pressure cast specimens with a diameter of 3 mm

13.7 References

and a length of about 25 mm increases for the gold alloy from about 24 lm to 91 lm and for the Al bronze from 30 lm to 170 lm when the mold temperature is increased from 1008C to 1000 8C. The casting result is also influenced by the surface structure, i.e. the porosity of the mold. A modification of a commercial, phosphate-bonded investment reduces the surface roughness from Ra = 1.13 lm to Ra = 0.44 lm compared to the commercial investment. Furthermore, the dimensional accuracy of the cast parts is determined by the expansion behavior of the investment which needs to be in compliance with the shrinking of the cast metal during solidification and cooling. Here, for bending bars with a nominal width of 200 lm, a deviation of only 1.5 lm between cast parts and patterns was achieved. The achievable flow lengths are determined by the casting material, the filling pressure and the mold temperature. An aspect ratio of 60 was reached for the gold base alloy Stabilor® G and a filling pressure of about 25 bar, even for a relatively low mold temperature of 700 8C. But Al-bronze also allows for aspect ratios which are sufficient for most applications within the collaborative research center. In summary, the investigations have demonstrated that microcasting is an ideal fabrication method for metal parts in microdimensions, e.g. for a demonstrator consisting of a gear unit and a turbine, in small and medium series production. A further advantage of microinvestment casting, i.e. the possibility of casting real three-dimensional microstructures even with undercuts, will be explored in the near future. In addition, for a higher level of automation, the development of a microchill-mold technique using permanent molds instead of lost molds may be promising.

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Morral, Cobalt Superalloys; Brussels: Centre d’Information de Cobalt, 1979. Biosil®, Instructions for Use; Hanau: Degussa Dental GmbH. H. H. Schulz, Dental-Labor 1976, 24, 339–344. H. Salmang, H. Scholze, Keramik; Berlin: Springer, 1986. Degussa Dental GmbH, Hanau, personal communications and information. H.-W. Gundlach, in Zahnärztliche Werkstoffe und ihre Verarbeitung, K. Eichner, H. F. Kappert (eds.); Heidelberg: Hüthig Verlag, , 1996, Vol. 1, Ch. 2. D. Low, M. V. Swain, J. Mater. Sci.: Mater. Med. 2000, 11, 399–405. L. Borchers, in: Zahnärztliche Werkstoffe und ihre Verarbeitung, K. Eichner, H. F. Kappert (eds.); Heidelberg: Hüthig Verlag, 1996, Vol. 1, Ch. 1. R. Carter, Gold Technol. 2001, (32), 7–18 G. Baumeister, S. Rath, R. Ruprecht, J. Hausselt, Proc. Mater. Week, 2001, Munich (CD-ROM) http://www.sfb499.de/ ASME/ANSI B46.1, Surface Texture (Surface Roughness, Waviness and Lay), New York: American Society of Mechanical Engineers (ASME), 1995. DIN EN ISO 4287. Geometrische Produktspezifikationen (GPS) – Oberflächenbeschaffenheit: Tastschnittverfahren – Benennung, Definition und Kenngrößen der Oberflächenbeschaffenheit; Release 1998-10. DIN 4768; 1990-05. Ermittlung der Rauheitskenngrößen Ra, Rz, Rmax mit elektrischen Tastschnittgeräten – Begriffe, Meßbedingungen; Release R. Papst, Ermittlung von mechanischen Kennwerten an Mikrozugproben aus Stabilor G® (Au58Ag23Cu12) unter Zuhilfenahme des Verfahrens der optischen Dehnungsmessung, Report of the Institute for Materials Research I, University of Karlsruhe (TH), Germany, Karlsruhe: Karlsruhe University, 2004, Ch. 5.1.2.

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14 Microelectroforming of Metals G. Schanz 1, K. Bade 2 Intstitute for Materials Research III (IMF III), Forschungszentrum Karlsruhe, Germany 2 Institute for Microstructuretechnology, Forschungszentrum Karlsruhe, Germany

1

Abstract Microstructured parts made of pure metals and alloys can be fabricated by the combination of lithography and electroplating. In a process commonly known as ‘primary structuring,’ a photoresist deposited on a conductive substrate is exposed to UV light or X-rays through suitable masks. After developing the photoresist layer, the resulting structures are electroplated, yielding metallic microstructures after stripping of the remaining photoresist. The process can be easily repeated, allowing a batch-type series production of metallic microparts. In this chapter, an alternative route for the fabrication of electroformed metallic microcomponents, the so-called ‘secondary structuring,’ is discussed. In this process, micropatterned polymer surfaces are electroplated, resulting in inverse metallic replicas. The process described focuses on microinjection molding of polymers (see Chapter 10) and on electroforming of Ni. The replication is illustrated in both its standard form for low aspect ratio parts and in its advanced state of development for high aspect ratio products. Electrochemical procedures are presented not only for Ni but also for Au and some technically relevant alloys. In addition, the simultaneous co-deposition of ceramic particles suspended in the electrolyte and some auxiliary techniques around the process including electroless metal deposition are briefly covered.

Keywords electrochemical deposition; microelectroforming; replication; nickel; metals; alloys

14.1 14.1.1 14.1.1.1 14.1.1.2

Electrochemical Metal Deposition

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Electrodeposition of Nickel 397 Electrolyte 397 Bath Analysis and Control 397

Advanced Micro and Nanosystems Vol. 4. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31493-8

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14.1.1.3 14.1.2 14.2 14.2.1 14.2.2 14.3 14.3.1 14.3.2 14.3.3 14.3.4 14.4 14.4.1 14.4.2 14.4.3 14.4.4 14.4.5 14.5 14.5.1 14.5.2 14.5.3 14.5.4 14.5.5 14.5.6 14.5.7 14.5.8 14.6 14.7 14.8

Device and Process at Laboratory Scale Electrodeposition of Gold 399 Deposition of Alloys and Dispersions

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Deposition of Alloys 400 Co-deposition of Ceramic Particles During Electroplating Replication of Microstructured Polymer Molds

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Mold Fabrication, Pretreatments and Quality Requirements 402 Principles and Limits of Non-directed Nickel Deposition 403 Procedure of Replication by Nondirected Electroplating 405 Results of Nondirected Electroplating 405 Process Perspectives for High Aspect Ratios

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Directed Metal Growth on Partly Conductive Templates 408 Insert Injection Molding 409 Relevant Properties of Polymers and Two-component Test Templates 410 Initial Phases of Nickel Deposition on Conductive Substrates 411 Replication of Two-component Templates by Electroplating 411 Auxiliary Process Steps

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Manufacturing of Microstructured Polymer Parts and Property Requirements 413 Template Cleaning 414 Microwave Plasma Treatment 414 Metal Deposition by Sputter Processes 414 Electroless Metal Deposition 415 Two-step Replication, Using Copper or Silicone Rubber 416 Post-treatment of Electroplated Replicas 416 Template Removal 417 Conclusion and Outlook Acknowledgments 418 References 419

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14.1 Electrochemical Metal Deposition The intention of the chapter is to present mainly process-oriented information. Concerning materials aspects, the following treatment remains less comprehensive. Thus, for example, Cu as a suitable material for certain technical applications, is not included at all (see, e.g., [1–3]), whereas the focus on Ni is due to both its technical relevance and its frequent use as standard for demonstration purposes. General information about the application of electroplating and electroforming in microtechnology and special effects arising there can be found elsewhere [4, 5].

14.1 Electrochemical Metal Deposition

14.1.1 Electrodeposition of Nickel 14.1.1.1 Electrolyte The established industrial processes based on electrolytes of sulfate (Watts type) or sulfamate type [6] have been adapted to microsystems technology (MST) applications. At the Forschungszentrum Karlsruhe (FzK), the sulfamate bath has been chosen owing to its potential for low internal stress nickel deposition, developed since 1983 with respect to electrolyte, bath analysis and maintenance and applied for replication of polymer molds [7, 8]. The main components are nickel sulfamate (72 g/l Ni) and boric acid (38 g/l), a pH of 3.5 is maintained by sulfamic acid addition and the usual working temperature is 52 8C. According to comprehensive investigations on bath performance, analysis and control, minor constituents were excluded and the surfactant (tenside) optimized. Important progress in bath development was the demonstration of performance and long-term stability of a special fluorinated surfactant (150 mg/l) [9]. This is essential for reliable wetting of microstructured substrates and defect-free metal deposition with minimum internal stress.

14.1.1.2 Bath Analysis and Control Purity requirements for process stability and product quality were defined according to photometric (Ni), polarographic (Ni, boric acid, surfactants), inversevoltammetric (trace metal impurities) and chromatographic (HPLC, sulfamate and surfactant decay) investigations. Typical currently applied procedures comprise ca. daily pH measurement and correction by addition of sulfamic acid and ca. weekly nickel and boric acid analysis, e.g. by redox titration and pH titration, respectively. The use of sulfur-depolarized nickel as anode material in the process of cathodic nickel deposition provides sufficient anodic nickel dissolution for compensation of bath depletion. During several months some eventual boric acid loss is replaced. The thermal stability of sulfamate is sufficient to prevent long-term decay to sulfate. The excellent stability of fluorinated surfactant helps to avoid enrichment of organic impurities [10]. Entrainment of these into the bath is minimized by adding best quality feed water to compensate for evaporation loss. A cleaning procedure for organics removal consists of oxidative treatment, filtration, active carbon adsorption and addition of fresh surfactant. Depending on charge passage, a carefully controlled nickel sulfamate bath could serve for up to more than 1 year.

14.1.1.3 Device and Process at Laboratory Scale The basic principles of electrochemical deposition (plating, galvanic deposition) are well known. Electrochemical nickel deposition consists of [Ni(H2O)6]2+ complex stripping and Ni2+ cation discharging during passage of the Helmholtz double layer. As a side reaction, hydrogen evolution takes place with proton con-

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sumption and is therefore connected with pH increase [6]. Precipitation (and possible co-deposition) of basic hydroxides would occur above pH 6. At free surfaces the boric acid buffer capacity of the sulfamate bath is easily high enough, but in small structural cavities the bath exchange can be much more restricted [11]. Consequently, a pH of 3.5, towards the lower end of the generally used range, is chosen and the current density is kept low for MST applications. Compared with some industrial processes, e.g. the restoration of huge worn metallic components by deposition of thick nickel layers, large-series polymer metallization or metal surface finishing, the necessary electroplating equipment for MST purposes is moderate. The main potential for the production of microstructured metallic parts via replication can be demonstrated even in a glass beaker of ca. 4 l volume. A DC voltage supply, connected to the conducting substrate (cathode) and the counter electrode (anode), both immersed in the electrolyte, provides the Ni2+ ion reduction and nickel deposition on the substrate without/with parallel Ni2+ replacement, according to the inert or soluble kind of anode. On the technical scale, more precise boundary conditions (e.g. temperature control, stable bath composition and pH, favorable current density distribution) are possible with more sophisticated commercial setups (e.g. HEGA Apparatebau, Straubenhardt-Conweiler, Germany), including customer-given features as used at FzK and described below. The typical central component is a ca. 40 l pool inserted into a desk, both of polypropylene (PP). The stand contains electrical systems for supply, control and operational safety and is provided with an exhaust system against aerosols. Different power supply units are available. From a rear subdivision of the bath container, electrolyte is continuously pumped through a particle filter of 0.2 lm pore size and fed into the main compartment at the bottom to avoid the formation of a stagnant sump. A fixed bath level is kept by flow over a weir back into the rear part, thus removing any floating dust. For bath heating, two heater cartridges are immersed and a thermo-resistance sensor for temperature controller input is positioned here. Automatic level control in the rear part is realized using levitation sensors for feed water control and off-normal level alarm. The water supply system in stand-by mode is connected via a particle filter, magnet valve and a two-stage rinsing basin, minimizing electrolyte loss. Feed water quality (18 MX cm) is maintained by two-stage tap water filtration, active carbon adsorption and ion-exchange cleaning. All parts in contact with the purified water and the electrolyte are made of pure PP, PTFE or PMMA to avoid re-contamination. For anode support, the bath is equipped with two contact bars; usually one titanium basket is used, filled with ‘rounds’ of sulfur-depolarized nickel and surrounded by a bag of PP cloth to retain coarser residues of the dissolving rounds. The substrate with protected contact wire is fixed, normally in a vertical direction, on a central PMMA plate, ca. 15 cm from the anode. This is hanging down from a cathode contact bar, which can be moved laterally with variable amplitude via a motor-driven eccentric to improve bath convection and to counteract current density asymmetry. According to form and size, the template substrate

14.1 Electrochemical Metal Deposition

is equipped with a frame and/or a screen in order to diminish preferential deposition at the rims. 14.1.2 Electrodeposition of Gold

Gold electrodeposition is mainly applied for the fabrication of x-ray masks, because of the high absorbance of x-ray photons [12, 13]. Owing to this function, a pure and pore-free (high-density) gold deposit is desired. Further, the pattern geometry has to be copied in a nearly perfect manner, because geometry losses in the first steps of pattern generation can lead to a loss of the function of the microcomponents [14–16]. The gold sulfite electrolyte yields deposits with the desired properties. The metal-carrying species is the complex anion disulfitoaurate[I], [Au(SO3)2]3–. Owing to the relatively low stability constant, a surplus of sulfite is necessary for operation. Further components can be buffering compounds and stabilizers. Commercial electrolytes often contain some additives such as As(III) or Sb(III) compounds, which influence the deposit properties (especially roughness, hardness and internal stresses). Chemical stability of the bath is maintained at pH values > 7 and the mentioned surplus of sulfite. The stability of the gold plating electrolyte may increase on addition of thiosulfate [17]. An inert anode (titanium/platinum) is used. Therefore, regular additions of gold complex are necessary to maintain the gold concentration. Compared with the nickel sulfamate process described above, the gold electrodeposition from sulfite electrolyte is much more sensitive to impurities and aging. DC currents

Fig. 14-1 Detail of a LIGA X-ray mask. Shown is a single Au microstructure on top of the Ti/Ti oxide membrane

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or pulse currents can be applied. For DC operation, the current density is between 0.2 and 0.5 A/dm2. Typically a thickness of up to 40 lm is deposited into the cavities of a resist pattern on a Ti/nanoporous Ti oxide layer (LIGA) or on to Cr/Au metallization layers. Larger thicknesses are accessible with pulse plating. In Fig. 14-1 a gold structure is shown, which is part of a LIGA x-ray mask applied in the fabrication of a mold insert.

14.2 Deposition of Alloys and Dispersions 14.2.1 Deposition of Alloys

In addition to elemental metal deposition, alloys can be deposited in high aspect ratio resist micropatterns. Several alloy electrodeposition processes are known from surface finishing, in which several micrometer thick layers are deposited on the base material [18, 19]. In high aspect ratio MST, much larger thicknesses in the range of several hundred micrometers are necessary. Most of the work at FzK has been focused on Ni–Fe deposition, because of the soft magnetic behavior of this material. However, the mechanical properties are also of interest, because the hardness is superior compared with nickel and also because of the hardness loss during thermal annealing [20–22]. Unfortunately this greater hardness is combined with a loss in ductility.

Fig. 14-2 Ni–Fe microcomponent

14.2 Deposition of Alloys and Dispersions

The deposition is performed with a sulfate electrolyte (NiSO4 and FeSO4) at 50 8C. The composition of the electrolyte (Ni2+/Fe2+ ratio) and the applied current density strongly influence the resulting alloy composition and the formation of composition gradients in the growth direction. For a constant composition, low current densities are applied, which corresponds to a longer processing time compared with the nickel process. In addition, a stress-reducing agent, e.g. saccharin, has to be used because in alloy deposition the stresses generated during growth are always larger than in elemental metal deposition. With that process it is possible to manufacture Ni–Fe microparts with thicknesses up to 300 lm and a composition of 80:20 wt.% (Fig. 14-2). Electroforming of Ni–Co which also shows increased hardness is commercially available. In Fig. 14-3 a comparison of the hardness loss during thermal annealing is shown for selected compositions of Ni–Fe, Ni–Co and Ni. The hardness maximum of Ni–Fe alloys is due to precipitates of intermetallic phase

Fig. 14-3 Hardness of Ni–Fe, Ni–Co and Ni vs. annealing temperature. Annealing time, 1 h

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in fine dispersion or to a change from the metastable as-deposited state to a more ordered state. Generally, alloy microelectroforming needs tighter process control compared with the case of metals, e.g. of the concentrations of metal species and the additives or their degradation products. 14.2.2 Co-deposition of Ceramic Particles During Electroplating

Suitable dispersions of ceramic particles in metallic matrices can have interesting technically relevant properties. Compared with alloy deposition, the potential of combining higher thermal stability against strength degradation, hardening without undue loss of electrical conductivity, improved friction or wear-related behavior, etc., could be advantageous for particular applications. Particle fraction, kind, size and even form may exert decisive influences under different conditions. Factors determining the potential of a co-deposition procedure are chemical compatibility of powder and electrolyte, persistence of particle or aggregate suspensions, efficiency of the compound with respect to particle fraction, distribution and binding forces. Parameters according to available theories are the hydraulic conditions close to the substrate (diffusion interface), a variety of attractive and repulsive forces between solvated ions, particles and the surface (particle dynamics) and the applied current density (metal deposition rate) [23]. The co-deposition of ceramic nanoparticles from suspension in metal electrolytes is currently under study at FzK. The suspensions of ceramic particles of oxide and nitride type in nickel sulfamate electrolyte mainly examined showed incomplete dispersion and limited stability, as expected for such concentrated solutions. Counteractions against particle agglomeration such as surfactant additions and mechanically assisted homogenization are being evaluated. According to first results, fairly promising particle suspensions in the electrolyte and dispersions in the deposited metal are obtained. Deviations from the behavior of micrometer-scale suspensions require also the development of analytical methods in advance of considering special applications [24].

14.3 Replication of Microstructured Polymer Molds 14.3.1 Mold Fabrication, Pretreatments and Quality Requirements

According to the considered principle of electroforming, polymer molds from a precise and economic fabrication process such as microinjection molding or hot embossing are required. As the following is restricted to injection molding, Chapter 10 can be referred to for basic principles of mold fabrication and references.

14.3 Replication of Microstructured Polymer Molds

Concerning material aspects, PMMA and various other polymers can be selected for use as molds in the electroplating process. The mold concept has to consider all steps, nuclei deposition and growth, layer formation and cavity filling. The range of pertinent mold properties differs substantially from quality aspects for direct application of polymer parts. The main requirement related to the use of polymer molds as a substrate for metal deposition is electrical conductivity at the surface. Alternatives to meet this cannot be intrinsically conductive polymers, but addition of conductive fillers, coating with a conductive surface layer or ‘electroless metal deposition’ on the chemically pretreated mold surface [25, 26]. Concerning the last option, conventional metal coating processes for polymers [27] can be used for orientation, e.g. treatments for sufficient polymer/metal adhesion during the plating process, but a minimum polymer surface roughening is indicated here in the interest of optimum replication quality. Sufficient chemical and mechanical template stability during plating replication is a trivial need. Some bending of PMMA molds was ascribed to residual internal stress and swelling tendency in a wet medium. Tempering of this or other template materials in a fixed position before use in plating replication could be considered. Substrate pretreatments and steps of special process variants are briefly covered in Section 14.5. 14.3.2 Principles and Limits of Non-directed Nickel Deposition

The convective and diffusive exchange of bath components and products defines a theoretical upper limit for a stationary rate of nickel deposition; for deep holes, not reached by convective exchange, the slowest diffusion step determines that limit. In practice, however, the surface topography of the conducting substrate determines the current density field, thus counteracting against the favorably high ‘throwing power’ of the sulfamate bath. A systematic study of nickel deposition in copper test substrates with slit geometry has shown that no deposition takes place at the slit base for aspect ratios > 10, obviously because the local electric field is not high enough to start the charge transfer across the surface [28]. On the other hand, higher current densities result at the front of slits or holes and preferential metal deposition at their orifice leads to premature clogging of those structures and enclosure of electrolyte-filled cavities (Fig. 144). The described process variant of non-directed metal deposition on the whole substrate surface is restricted to aspect ratios < 10, if the mentioned cavities can be tolerated with respect to their mechanical and thermal consequences; if not, a stronger restriction to an aspect ratio (AR) of 5 is indicated. The coverage of the complete substrate front, in terms of the metallic product the automatic formation of a continuous base plate, is favorable for an intended production of copies of injection molding inserts (Fig. 14-5; see also [29]). In contrast, it is unfavorable for batch fabrication of isolated microparts, since this requires a base plate removal step for separation of the array. It is mentioned for comparison that an electroless deposition of, e.g., Ni(P) is not accompanied by the described topography-related nonhomogeneous current

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Fig. 14-4 Scheme of mold, start layer and cavity formation during electrodeposition of metal

Fig. 14-5 Carbon black-filled PA 12 mold and nickel replica with as-grown contours

density; since no electric field is applied, high-AR structures can be filled more completely [28]. For the electroplating step of the LIGA process, the described field distribution is not relevant and filling of nonconducting cavities with conducting base should be possible even for high ARs. An alternative procedure for obtaining partially metalized mold patterns is the deposition of metal clusters, which is possible in a strictly directed way in order to exclude the coverage of mold side walls [30]. Convective diffusion and current density distribution are also important parameters in the electrodeposition step of all those methods [31, 32].

14.3 Replication of Microstructured Polymer Molds

14.3.3 Procedure of Replication by Nondirected Electroplating

According to the FzK standard process for nickel deposition [33], the pretreated mold (see Section 14.5) is contacted to copper wire, insulated at its rim, surrounded by a frame and mounted on a support plate. Substrate prewetting by electrolyte or surfactant solution and supporting evacuation pulses remove any caught air bubbles. A delay of a few minutes after insertion into the bath is sufficient for equilibration of concentrations and temperatures. Usually the current-controlled deposition option is chosen and a limiting voltage is set. For optimum filling of the structural cavities, a current density of *0.5–1 A/dm2 is applied, corresponding to a deposition rate of *6–12 lm/h. After completion of filling of the structures, the current density can be increased to 2 or even 4 A/dm2, but this favors nodular metal deposition [34]. An improved distribution of the resulting macroscopic current density can be achieved by using a frame of previously determined size, contour and distance. 14.3.4 Results of Nondirected Electroplating

The section is structured according to several aspects of product quality, as demonstrated for FzK standard process steps or applied alternative steps, giving more direct evidence [33]. The electrochemical replication step of the described basic process has not shown any fundamental problems in practice. On the one hand, careful procedures give reliable results; on the other hand, the limitations related to formation of cavities in high aspect ratio replicas cannot be overcome. In the range 1–10 A/dm2, current efficiencies above 98% were determined, and down to the usually applied much lower current densities the deposit microstructures are fairly similar. Porosity due to adhering hydrogen bubbles is avoided by having a sufficient surfactant content. Formation of ‘trays’ at the rim of the product, related to gas bubble capture by the frame, is prevented by intensive bath circulation. With increasing current density and process duration, the free metal surface becomes scalloped, especially at the rims and above spikes of the mold. Preferential growth of columnar crystal aggregates can result in recessed grain boundaries between them, which can also trap gas bubbles. Excellent resolution of the replication in nickel was demonstrated for the example of a holographically exposed and chemically developed resist layer on a glass substrate with a regular matrix-like arrangement of sinusoidal surface waves. The SEM image of the nickel replica demonstrates nanometer-scale resolution and absence of replication defects (Fig. 14-6) [35, 36]. Preservation of the macroscopic shape of the polymer mold by the nickel replica cannot be determined directly, since the mold itself can be much more susceptible, e.g., to bending or torsion. Unambiguous results were obtained by nickel deposition on thick and carefully flattened copper substrates, finally stripped by selective dissolution. The relaxed nickel foils indicated minor bending in the

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Fig. 14-6 Resist with array structure (back) (courtesy FhG ISE, Freiburg, Germany) and highly resolved nickel replica (front), SEM

sense of tensile internal stress (concave growth front, convex structure side). The trend towards bending increases with preferential rim deposition and during deposit thickening under high applied current density, by which tool production was simulated. Templates of silicone rubber, produced by replication of ‘master parts,’ were applied for repeated secondary replication in metal, to demonstrate the reproducibility of the replication step with respect to eventual artifacts, dimensional deviations and bending. Successful reproduction, absence of structural artifacts, fair dimensional stability and acceptable deviation from flatness were identified, in spite of the mechanically weak mold. In addition to light microscope measurement protocols, topographic measurements were performed by evaluating line and area scan data from a light-optical contour and depth detection device (‘microglider,’ Fries Research Technology, Bensberg, Germany). The software allowed two- and three-dimensional illustration of the topography and extraction of profiles and dimensional data. For example, a nickel replica of a silicone rubber mold, this mold itself and the original, a deep-etch structured silicon part of 60 ´ 20 mm, were compared. The nickel product showed a convex structure side with total bending of 70 and 50 lm in the longitudinal and lateral direction, respectively, both referred to a virtual flat base plane. This amount of bending could not be tolerated for application as a tool without compensation, but would be less important for the elements of the separated array. Dimensional product qualification can be sophisticated and should be restricted to the essential data for a given application. A comparative study included an Ni–Co original (UV lithography), the (inverted) silicone rubber replica and the daughter copy in Ni. Another study used master parts of silicon with test structures of EPON SU-8 UV resist, replicated in one or two steps. The determined relative length deviations of two-step daughter copy from origi-

14.3 Replication of Microstructured Polymer Molds

Fig. 14-7 Array of nickel test columns during the mold dissolution step, SEM

Fig. 14-8 Nickel parts on adhesive tape, after mechanical array separation, SEM

nal were below 1% and the length dimensions of replicated structures of the same type were reproducible to <0.3%. Acceptable dimensional precision can thus be maintained even for two-step reproduction via the mechanically unstable silicone rubber replication step which was used for demonstration purposes. The final step of array separation was demonstrated for elements of nickel test structures (Figs. 14-7–14-9) under laboratory conditions. Its adaptation to an industrial batch manufacturing procedure will have to decide between the op-

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Fig. 14-9 Illustration of series production potential for columnar test parts

tions of micropart production as a bulk of separate pieces or via an array magazine, containing removable parts.

14.4 Process Perspectives for High Aspect Ratios

The previous section was focused on a ‘standard process’ of electroforming by polymer mold replication, which is of basic kind in so far as it is only successful for the fabrication of metallic parts with structure aspect ratios < 5. This limit was shown to be the inevitable consequence of non-directed metal growth within fully conductive structure cavities, which cannot be completely filled owing to clogging of their orifice. Now, an advanced concept under development at FzK is presented by which this limitation can be avoided [37]. 14.4.1 Directed Metal Growth on Partly Conductive Templates

The principle is well known, e.g. from the electroplating step of the LIGA process. The exposed areas of a conductive substrate, from which the nonconductive resist layer is removed during resist development, allow metal deposition, whereas the remaining resist defines the side-wall contour of the metal micropart. Some alternative principles by which strictly directed metal deposition in a structured ‘mold’ with inert side contours can be obtained are given in [38]. Transcribing this principle to polymer mold replication requires molds with conductive base and nonconductive side walls and a suitable process for their fabrication (Fig. 14-10).

14.4 Process Perspectives for High Aspect Ratios

Fig. 14-10 Principles of electrodeposition on microstructured polymer molds

‘Selective’ metal deposition was considered for the first time at FzK, when microinjection molding of carbon black filled polymer was found to give higher surface conductivity at the base plate of the product than at the sides and fronts of small protruding structures [25]. Studies on their replication showed that indeed the base plate was quickly covered by metal whereas the micropatterns were overgrown more slowly beginning from the base. However, the resistance gradient was not sufficient to avoid the finally complete metal coverage of the structures by which the mold conductivity field lost its influence. The current concept for quasi-infinite conductivity gradients at the mold surface is to combine a nonconductive and a conductive polymer in a compound in which the nonconductive polymer structures protrude from a conductive base. The technical process to realize this is under development for dimensions relevant for MST application of the two-component polymer compounds and their use as sacrificial templates for electroforming of metallic parts (see Chapters 10 and 11 for comparison). 14.4.2 Insert Injection Molding

In a first polymer injection molding process step, a conducting polymer is injected into a first tool insert, structured for base plate formation. This part is inserted into a second microstructured tool into which nonconductive polymer is injected in the second molding step (Fig. 14-11). The two polymers must be compatible with respect to thermal/mechanical stability of the first during the second step and adhesion of the second polymer. With respect to compound cohesion, it is especially important to avoid wedge-shaped slits at free interface contours of the compound which would cause artifacts in the electroforming process. Together with the choice of conducting filler, many materials-related aspects have to be investigated and judged for relevance. Together with the complex technical require-

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Fig. 14-11 Process scheme of two-component injection molding

ments, tool design and fabrication, flow and temperature distribution modeling, the development is too challenging to address its present state here in further detail. Some more information can be found in Chapter 10. 14.4.3 Relevant Properties of Polymers and Two-component Test Templates

As commercially available raw materials the polymers PMMA, POM, PA 12, PA 6.6, PS and PEEK were processed to flat plates or parts with test micropatterns. Carbon black filler ‘C’ and carbon fibers filler ‘CF’ were present in variants, here denoted PA 12 (C), POM (C), PA 6.6 (CF) and PEEK (CF). In contrast to the intrinsic conductivity of special polymers not applicable here, the filler-related conductivity of bulk and surface may depend on filler content, particle or fiber size and distribution. Especially the surface conductivity, measured by the two-point and four-point principle, is relevant for the intended metal deposition. Owing to the mass fractions (15–20%) of carbon black, which were sufficient to reach the percolation limit of conductivity, and the small particle size (<100 nm), the measured surface resistance was of the order of 10–100 X/cm for all C-filled polymers and fairly homogeneous. The surface resistance of PA 6.6 (CF) was of the same order, that of PEEK (CF) was about one order of magnitude higher and the results obtained for both were more scattered and unstable. This may be the consequence of a more isolated fiber distribution or less frequent penetration to the surface, in spite of rather high volume fraction. The distribution of the long fibers (*100 lm), which are oriented mainly parallel to the flow direction of the injected mass, is not as favorable for resistance homogeneity as a frequent occurrence of fiber crossovers would be. It is mentioned that the measured more inhomogeneous surface conductivity of ‘CF compounds’ is reflected in the SEM investigation of specimens: The distribution of induced charging contributed to their image contrast.

14.4 Process Perspectives for High Aspect Ratios

14.4.4 Initial Phases of Nickel Deposition on Conductive Substrates

Common for all kinds of conductive substrates, the nucleation phase is that of metal deposition at dispersed sites, favorable for charge transfer. Next, growth of the nuclei and finally the formation of a closed cover layer follow. The initial phases of nickel deposition on various filled polymers were studied and compared. Carbon black-filled substrates showed fairly fast formation of a closed and fine-grained nickel layer, similar to the deposition on metallic substrates. Starting from preferred spots, the metal cover was closed during their growth. In contrast, carbon fiber-filled polymers remained on average much longer in the phase of partial metal coverage, owing to an insufficient number and unfavorable distribution of nuclei. Insufficiently conductive spots of the polymer surface were not ‘competitive’ enough to avoid the resulting current density distribution, which favored the already covered areas further. A complete layer was only reached owing to lateral growth of earlier overgrown areas. In a schematic way, hemispherical growth of an array of nuclei can illustrate the more complex real situation. Accordingly, the growth of those hemispheres would have to continue until they came into mutual contact. The delayed formation of a closed metal layer on carbon fiber-filled polymers can be tolerated, provided that the longer direct electrolyte exposure of the substrate and the thickness variation of the closing layer play no role. The compact description of the comprehensive test program on conductive polymer start phase replication is concluded by the remark that the polymer/filler combination determines the observed range of nickel deposition characteristics, whereas the parameters of the plating process, mainly the applied current density, cannot be decisively optimized. In a crude simplification, carbon black filler is to be preferred to carbon fiber filler. 14.4.5 Replication of Two-component Templates by Electroplating

In a study on cavity filling by oriented nickel growth, two-component templates were simulated by test structures of slit and hole geometries, metal base and nonconductive side walls. Some problems with incomplete electrolyte wetting were solved and the current density had to be set low enough for diffusive mass transfer. Continuous cavity filling was registered during the long-term test exposures. Some influence of convective electrolyte exchange in slits was registered, leading to asymmetric growth front contours in several of the systematic tests. The second template component, the nonconductive polymer, determines the side contour of the nickel replica structures and is therefore most important with respect to dimensional stability, precision and surface quality. Excellent replication in nickel can be expected for any polymer, favored for advanced twocomponent template production, thus reproducing the precision of the mold. Consequently, the ‘filling phase’ of the electroplating process can be seen as a qualification test for the polymer template.

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Two-component templates from insert injection molding pilot production campaigns allowed the study of different polymer combinations and test structure designs. The latter can be classified into an ‘island type‘, small individual resist structures on a large conductive base, and ‘hole type‘, continuous resist cover on the conductive base which is not covered only at distributed small structure cavities. The corresponding nickel replicas are microstructured plates or isolated microparts illustrative for tool formation and series production of parts, respectively. Combinations of the two basic structure types also have technical relevance, but are not considered here. The replication of ‘island-type’ templates allows free electrolyte access and requires no stringent current density restriction. With the example of close structures of rosette and cross form, it was verified that the filling of the narrow space between them takes place with comparable growth rate (Figs. 14-12 and 14-13). Reliable replication of templates of ‘hole type’ depends on a sufficient local conductivity of the base plate at the positions of the structure cavities. Theoretically, at least one nucleus per cavity has to grow in order to start metal deposition at the base, which will finally be overgrown within each cavity even in this worst case. However, for geometric reasons, mass transfer restrictions could result in replication artifacts at corners: Rounding of edges would indicate the distance of necessary lateral nuclei growth. Replication of ‘hole-type’ templates was successfully demonstrated: by oriented nickel growth, the basic principle of the development, the massive filling of narrow cavities is no problem. Cavities of conductive polymer (one-component template) and similar form have shown the typical three-dimensional columnar grain growth and void formation below the cavity annulus. Filling of two-component templates was successful, e.g. for

Fig. 14-12 Front view of Ni deposit growth on PA 6.6 (C), replicating POM structures

14.5 Auxiliary Process Steps

Fig. 14-13 Ni replica of PA 6.6 (C) + POM template, SEM, view on base

an aspect ratio of ca. 40, a calculated value based on a 1 mm structural height and 25 lm minimum lateral width [38–40].

14.5 Auxiliary Process Steps 14.5.1 Manufacturing of Microstructured Polymer Parts and Property Requirements

Basic principles of microinjection molding are given in Chapter 10; the alternative hot embossing is not covered. The polymers considered include PMMA, intensively studied for reference purposes, other processed amorphous and partly crystalline thermoplastic materials and polymer compounds filled with carbon black powder or carbon fibers. The application of polymers for structure replication by electroforming requires a property range which is different from that suitable for their direct use as a product. Important factors are dimensional stability during the replication process, chemical stability in contact with the electrolyte, sufficient electrical conductivity through the bulk or at the surface, alternatively compatibility with a process step for metal coating, etc. Details are given for the aspect of dimensional stability of PMMA, as an example for manifold requirements. In the injection molding process, plasticized PMMA is injected at ca. 160–170 8C into the cavity at ca. 120–140 8C tool temperature. To achieve short cycle periods, forced cooling to 80–90 8C takes place under sustained pressure until the part is removed. Internal stress relaxes to a residual level given by final cooling. Tempering of such a part in air at the electrolyte working tem-

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perature (52 8C) resulted in considerable bending to a convex shape of the structured side. Even greater bending after exposure to hot water or electrolyte indicated swelling due to absorption of some water and shrinking with its delayed release during drying. The tests have shown that mold contour stability is not a trivially given fact. 14.5.2 Template Cleaning

When cleaning the polymer parts to remove surface contaminants, possible migration of intrinsic additives, added to support the demolding step, has to be considered. Some commercial PMMA granulate seemed to contain an additive on a fatty acid base which continued to migrate to the surface of parts during their storage; rinsing with alcohol alone was not sufficient for complete removal, which was achieved in the next standard preparation step. 14.5.3 Microwave Plasma Treatment

A versatile vacuum facility, allowing consecutive substrate conditioning in a microwave plasma (2.45 GHz) and the sputter deposition of metals (DC planar magnetron), was used for the development of the physical process steps for the preparation of polymer templates. Microwave plasma treatment, studied under a range of conditions (Ar, Ar + O2, Ar + NH3; 1–10 Pa; < 1 kW) proved to be very effective in increasing the adhesion of the subsequently sputter-deposited layers of different metals on various polymer surfaces. Conditions should comprise simple process control, access to recessed surfaces, applicability to both delicate and chemically stable substrate materials and tolerable surface roughening due to a substrate etching rate of the order of only 10 nm/min. For PMMA, electron spectroscopy for chemical analysis (ESCA) analyses, contact angle measurements (Kruess G29, 3 ll, ethylene glycol/water test fluids) and their evaluation by the method according to Owens, Wendt, Rabel and Kaelble (OWRK) identified the resulting chemical surface modification and the considerable increase in the polar fraction of the surface energy. In the interest of a simple standardized step, the microwave plasma treatment can be initiated below 5 ´ 10–2 Pa, performed in argon purge gas of (3–9) ´ 10–2 Pa for 5 min and applied to different polymer types [33]. 14.5.4 Metal Deposition by Sputter Processes

This metallization method can be seen as the default option within the ‘basic standard process’ of electroforming developed at FzK. Since PMMA was not available as a conductive compound, the necessary metal deposition step was developed with respect to the combination of sufficient surface conductivity and

14.5 Auxiliary Process Steps

adhesion. A test program covered the target materials Au, Ag, Cu, Cr and Ni and also Ni(P) and Pd with respect to the electroless nickel (EN) deposition option. The parameters and their ranges were pressure 1–0.1 Pa, substrate distance 300–100 mm, duration 10 s–30 min, current 14–500 mA and corresponding power 5–200 W. Layer growth rates were determined as function of the parameter set and served for orientation of the limits of tolerable energy deposition rate. Optimum adhesion to various polymer substrates was achieved for consecutive sputtering of Cr and Cu from two targets, to avoid interruption by air inlet [33]. In the interest of a simple process step, it was proved to be sufficient to sputter a layer of ca. 100 nm of gold from a target/substrate distance of ca. 80– 90 mm, at 0.8 Pa pressure and 36 W sputter power (corresponding to 100 mA), which requires a net duration of only 1.5 min. Insufficient coverage of the sides and base of recessed template structures on the one hand and decreasing adhesion and resolution for excessively thick layers on the other define a broad range of uncritical deposit thickness. The application to structures with demanding aspect ratio can be adapted by reducing the substrate distance to 30– 40 mm, thus improving the penetration. Then the treatment has to be limited to 10 s at 200 W power (500 mA) or the power must be reduced. Hence, cavities of 200 lm lateral and 1 mm vertical dimension (AR 5), for example, can be metal coated. 14.5.5 Electroless Metal Deposition

For more demanding aspect ratios, for which the sputter deposition technique alone would fail to provide sufficient galvanic start conductivity, a chemical process of wet template treatment for electroless nickel deposition can be considered, following physical ‘seed’ (technical term for catalytically active nuclei) metal particle deposition in a previous Ni(P), Ni or Pd sputter step. Alternatively, surface etching, seed crystal deposition and metal layer growth can be performed via wet process steps alone. Electroless metal deposition (synonym: chemical deposition) is a process supported by charge exchange between different bath components. Generally, positive metal ions are reduced to metal, which is deposited on a specially prepared surface, parallel to the oxidation of a reducing agent in the bath volume. This requires avoiding ‘parasitic’ metal seed precipitation in the volume by the presence of a stabilizing agent as third component. Surface preparation has to provide seed metal crystals on a polymer surface which are able to initiate the chemical metal deposition. This preparation step may include exchange of noneffective primary seed by cementation into effective secondary seed. Among the developed electroless processes, those for nickel, copper and gold are usually applied. Electroless nickel (EN) process variants [41], commercially available, differ in deposition rate according to process temperatures up to 95 8C and in the phos-

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phorus content from side reaction of oxidation of hypophosphite to phosphite. Polymer- or copper-coated substrates can be Ni(P) coated after seed deposition treatment with solutions containing Pd in ionic or colloidal form. It is mentioned that the high internal stress which is built up generally in an EN deposit increases the demand for sufficient polymer surface adhesion forces compared with galvanic Ni deposition [27]. Process stability requires bath composition and process parameter control, in addition to frequent or continuous component dosage and pH adjustment. These increased efforts are justified if no alternative option is available or in the case of application of the EN layer as a functional surface of the nickel product, which is formed by galvanic back filling in a subsequent electroplating process step. Use of Ni(P) as a cover layer is an option with respect to applications requiring improved wear-resistant and/or corrosionresistant surfaces. Formation of three-dimensional Ni(P) parts is generally not considered owing to the very limited ductility of the alloy. 14.5.6 Two-step Replication, Using Copper or Silicone Rubber

A technique for double structure inversion and thus reproduction of an original part can be needed for various purposes: a mechanically weak part must be converted to a tool, a valuable original is to be reproduced non-destructively, a structuring process is successful only for the negative contour or an inadequate material, etc. One of the briefly mentioned methods is structure replication by copper electroplating as the first step, followed by (destructive) removal of the original part. This provides a mold for the second step, structure replication in nickel. Selective copper dissolution versus nickel is used for destructive mold removal from the nickel product part. Replication of an original part by embedding into silicone rubber resin and separation after resin curing offers the chance to retain the original, owing to the outstanding plasticity of silicone rubber and the weak adherence to most substrate materials. The second replication step requires silicone rubber metallization and contacting for structure replication in nickel. Precautions mainly against replication artifacts related to silicone rubber deformation are essential in this procedure. 14.5.7 Post-treatment of Electroplated Replicas

Mechanical or spark erosion techniques can be chosen according to the dimensions of the electroformed raw product and the contours finally needed. Preferentially, the mold remains as part of the treated compound as long as possible in order to protect the structures. Options are to convert the irregular front (free surface) of the raw part to a plane ‘base plate’ (daughter tool fabrication) or to a ‘frame’, allowing interim handling and storage of the thus fixed array elements for their final removal, e.g. by laser cutting and mounting. The latter includes

14.6 Conclusion and Outlook

the option to separate the elements, but to retain them fixed to individual frames for further handling. For positioning and orientation, all treatments require markers at the rim of the structured field, which have to be provided within the structure design step. An alternative procedure is indicated if the application requires precise structural height of the product or product array. In this case, the first step can be the removal of the mold base plate combined with the front plane leveling of the embedded structures. Next, complete product base plate removal and height calibration of the slab would follow. Consequently, the removal of the mold material of the slab would result in separate array elements, which might be tedious to handle. This problem is simple to avoid by adapting any kind of interim fixed positioning compatible with the mold removal step and the processing line for microsystem mounting. 14.5.8 Template Removal

PMMA template removal from the metallic product is easily effected by dissolution in acetone or ethyl acetate. Careful rinsing is necessary; solvent recovery by distillation can be part of an industrial process, as well as ultrasonic process acceleration. Other polymers may require special organic solvents or alkaline solutions for which nickel stability is provided. For completeness, it is mentioned that selective EPON SU-8 removal from nickel is challenging [42], whereas deep-etched silicon substrates can be destructively removed in alkaline solutions and copper or brass in a bath based on carbaminate [43]. Silicone rubber curing, e.g. in contact with PMMA, leads to chemical interactions which exclude this material combination, whereas the silicone rubber contact forces are generally low enough to allow its removal by tension or bending.

14.6 Conclusion and Outlook

Economic processes based on electrochemical deposition are being developed for the industrial series fabrication of microstructured parts or components in metals and alloys. One concept is the batch and pilot series fabrication of microstructured metallic parts via the LIGA process sequence. Metal and alloy electrodeposition is briefly described and the potential is illustrated. Investigations on the co-deposition of ceramic sub-micrometer particles are mentioned. The concept which is mainly considered here is to use microstructured parts from polymers as sacrificial templates for replication of their structure. The availability of such molds from injection molding or hot embossing processes, which are developed in parallel towards microscale dimensions and high aspect ratios, is assumed. The initial step within this concept was to identify and characterize a ‘basic standard’ process for nickel deposition by electroforming using injection-molded

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PMMA templates, restricted to complete structure cavity coverage and non-directed filling. In a physical two-stage pretreatment, the PMMA mold is cleaned and preconditioned for metal adhesion by microwave plasma exposure and metal coated by gold sputter deposition. The nickel replica of the polymer mold is built up in a nickel sulfamate–boric acid electrolyte. The raw replica is spark eroded for copied mold insert production or milled to the required structural height from the base plate or for complete removal of the latter. The process is completed by PMMA dissolution and product quality control. Restriction of the process to a minimum number of reliable and compatible process steps is favorable for economic series production. General quality requirements are met according to microscopic inspection and dimensional measurement of structural details for parts of various test designs. Pilot series fabrication would provide conclusive information on product quality and reproducibility versus process parameter tolerances. The described basic process is suitable only for aspect ratios < 5. This is due to the inevitable facts that the gold sputter deposition rate for higher AR cavities is insufficient and that their filling in the plating step is not possible without void formation. An advanced process for the production of high aspect ratio metal microparts via polymer templates is under development. The task depends on the successful development of template fabrication in an injection molding process. A suitable template consists of a conducting base and nonconducting structures so that the directed metal growth from the base in the structure replication by electroplating is not aspect ratio limited. Promising results have been obtained using insert injection molding (see Chapter 10), e.g. successful filling of twocomponent templates was demonstrated for a calculated aspect ratio of ca. 40 and 25 lm lateral width. Templates with different test structures were replicated in nickel, thus precisely transferring their form and surface quality. The insert injection molding process itself will be further developed with respect to small dimensions, materials (optimized polymer combinations) and technical aspects (design and injection parameters). In addition the two-component injection molding process will be developed. Electrochemical replication, for which no directly related problems are foreseen, will profit from this development and it will also provide the necessary feedback for template fabrication. The full potential of the sophisticated combined process remains to be demonstrated.

14.7 Acknowledgments

Phases of the process development were supported by the Deutsche Forschungsgemeinschaft (DFG) under Program No. SPP 1012. The contributions of N. Holstein and J. Lorenz are gratefully acknowledged. We also thank C. Nold, B. Kloppe, K. Naumann, G. Finnah, E. Oskotski and V. Piotter for support and cooperation.

14.8 References

14.8 References 1 J. Elze, Metall 1954, 8, 519–526; 1955, 9, 2

3

4 5 6

7 8

9

10 11

12

13

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104–109. P. C. Andricacos, C. Uzoh, J. O. Dukovic, J. Horkans, H. Deligianni, IBM J. Res. Dev. 1998, 567–574. T. Daenen, W. Messing, M. v. d. Hout, E. J. v. Loenen, B. v. d. Bossche, Plating Surf. Finish. 1999, 78–82. L. T. Romankiw, Electrochim. Acta 1997, 45, 2985–3005. M. Datta, D. Landolt, Electrochim. Acta 2000, 45, 2535–2558. S. A. Watson, Nickel Electroplating Solutions, NiDI Technical Series No.10047; Nickel Development Institute, 1989. S. Harsch, W. Ehrfeld, A. Maner, KfK 4455; 1988. W. Bacher, K. Bade, K. Leyendecker, W. Menz, W. Stark, A. Thommes, in: Electrochemical Technology; Kodansha, 1996, p. 159. W. Stark, M. Saumer, B. Matthis, Galvanotechnik 1995, 86, 3931–3935; 1996, 87, 1107–1111. K. Bade, A. Fath, W. Leskopf, Metalloberfläche 1999, 53, 34–38. K. Leyendecker, W. Bacher, K. Bade, W. Stark, Thesis (K. Leyendecker), FZKA 5594; 1995. W. Bacher, P. Bley, H. Hein, U. Klein, J. Mohr, W. K. Schomburg, R. L. Wang, Y. M. Desta, R. K. Fettig, J. Goettert, H. Hein, P. Jakobs, J. Schulz, J. Micromech. Microeng. 2004, 14, 722–726. W. Bacher, P. Bley, H. Hein, J. Mohr, W. K. Schomburg, R. Schwarz, W. Stark, KfK Nachr. 1991, 23, 76–83. W. Chu, M. L. Schattenburg„ H. I. Smith, Microelectron. Eng. 1992, 17, 223–226. J. J. Kelly, N. Yang, T. Headley, J. Hachman, J. Electrochem. Soc. 2003, 150, C445–C450. T. Ogawa, T. Soga, Y. Maruyama, H. Oizumi, K. Mochiji, J. Vac. Sci. Technol. B 1992, 10, 1193–1196. T. A. Green, M. J. Liew, S. Roy, J. Electrochem. Soc. 2003, 150, C104–C110. D. Landolt, Electrochim. Acta 1994, 39, 1075–1090.

19 B. Loechel, S. Jung, H. Zimer, H. M.

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24 25

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27

28 29 30

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32 33

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Heuck, J. Chlebek, Electrochem. Soc. Proc. 1999, 33, 307–316. A. Fath, W. Leskopf, K. Bade, W. Bacher, in: “Microfabricated Systems and MEMS IV,” P. J. Hesketh, S. S. Ang, W. E. Bailey, J. L. Davidson, H. H. Hughes, D. Misra (eds.), PV 2000-19; Pennington: Electrochemical Society, 2000, p. 7. C. Cheung, F. Djuanda, U. Erb, G. Palumbo, NanoStruct. Mater. 1995, 5, 513– 523. M. Thuvander, M. Abraham, A. Cerezo, G. D. W. Smith, Mater. Sci. Technol. 2001, 17, 961–970. J. L. Stojak, J. Fransaer, J. B. Talbot, in Advances in Electrochemical Science and Engineering, R. C. Alkire, D. M. Kolb (eds.); New York: Wiley-Interscience, 2001, Vol. 7, Chapter 4. J. Lorenz, unpublished results, 2004. V. Piotter, J. Hausselt, J. Konys, R. Ruprecht, in: Proc. MicroSystem Technologies ‘96, Potsdam, 17–19 September 1996; 1996, pp. 193–198. A. Thies, J. Konys, J. Hausselt, in: 35. Jahrestagung der DGO, Lübeck-Travemünde, 24–26 September 1997; 1997. W. Riedel, Funktionelle Chemische Vernickelung, 1st edn.; Saulgau: Eugen G. Leuze, 1989. A. Thies, G. Schanz, E. Walch, J. Konys, Electrochim. Acta 1997, 42, 3033–3040. U. Michelsen-Mohammadein, Metalloberfläche 1998, 856–858. A. Thies, V. Piotter, J. H. Hausselt, O. F. Hagena, Microsyst. Technol. 1998, 4, 110– 112. G. Engelmann, R. Engelke, M. Heinrich, G. Grützner, J. Wolf, H. Reichel, in: Proc. MicroSystem Technologies, Munich, 7–8 October 2003, H. Reichl (ed.); 2003, pp. 265–272. A. Kondomitsos, H. Löwe, K. Russow, Galvanotechnik 2001, 16–21. G. Schanz, E. Walch, M. Guttmann, C. Nold, J. Konys, Internal Report; Forschungszentrum Karlsruhe, 2001. A. Thies, G. Schanz, E. Walch, J. Konys, J. Hausselt, in: 3. Statuskolloquium des

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35

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PMT, 2.–3 April 1998; FZKA 6080, 1998, pp. 109–114. W. Bacher, K. Bade, J. Konys, G. Schanz, Nachrichten – Forschungszentrum Karlsruhe 1998, 30, No. 3–4, 251–260. K. Bade, A. Fath, M. Guttmann, E. Walch, in: 4. Statuskolloquium des PMT; FZKA 6423, 2000, pp. 191–192. M. Guttmann, E. Walch, G. Schanz, J. Konys, Metalloberfläche 2000, 54, No. 11, 16–28 (Part I); 2001, 55, No. 1, 34–36 (Part II). N. Holstein, V. Piotter, R. Ruprecht, G. Schanz, in: Proc. MicroSystem Technolo-

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gies, Munich, 7–8 October 2003, H. Reichl (ed.); 2003, pp. 273–281. N. Holstein, V. Piotter, J. Lorenz, E. Oskotski, R. Ruprecht, G. Schanz, J. Hausselt, Adv. Eng. Mater. 2004, 6, 643–653. N. Holstein, G. Schanz, J. Konys, V. Piotter, R. Ruprecht, Microsyst. Technol. 2005, 11, 179–185. K. Stallmann, Galvanotechnik 1984, 75, 142–149; 1251–1260. H. Lorenz, M. Despont, P. Vettiger, P. Renaud, Microsyst. Technol. 1998, 4, 143– 146. W. Bacher, R. Schwarz, Patent Germany DE-PS 4024909; 1993.

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15 Further Ceramic Replication Techniques H.-J. Ritzhaupt-Kleissl 1), H. von Both 2), M. Dauscher 1,2), R. Knitter 1) Institut for Materials Research III (IMF III), Research Center Karlsruhe, Germany 2) Institute of Microsystem Technology, University of Freiburg, Germany

1)

Abstract This chapter gives a brief review of replication techniques, which are not part of the SFB 499 but which are appropriate for the fabrication of microstructured ceramic parts or ceramic microparts. These are the casting and embossing of green tapes, electrophoretic deposition, gel casting, direct coagulation casting and the variants of soft molding.

Keywords tape casting; embossing; electrophoretic deposition; microstructured electrodes; membrane method; gel casting; direct coagulation casting; soft molding; replica molding; microtransfer molding; micromolding in capillaries; solvent-assisted micromolding 15.1 15.2 15.3

Introduction 422 Embossing of Ceramic Tapes 422 Electrophoretic Deposition 426

15.3.1 15.3.2 15.3.3 15.3.4

Principle 426 Suspension Stability 427 Kinetics of Electrophoretic Deposition 428 Fabrication of Microstructured Components by Electrophoretic Deposition 429 Sol–Gel Electrophoretic Deposition of Microstructured Ceramics 432

15.3.5 15.4 15.5 15.6 15.6.1 15.6.2 15.6.3 15.7 15.8

Gel Casting 432 Direct Coagulation Casting (DCC) Soft Molding 437

436

The Soft Molding Technique and Its Variants 437 Soft Molding of Preceramic Polymers or Precursors 439 Soft Molding of Ceramic Suspensions 441 Conclusion 444 References 444

Advanced Micro and Nanosystems Vol. 4. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31493-8

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15.1 Introduction

In addition to the ceramic replication techniques investigated in SFB 499 and described in previous chapters, some further replication techniques will be referred to here. These techniques have in common that ceramic replicas are generated from molds, whereas the mold materials and the methods by which the molds or the original masterpieces are fabricated may be different. For the ceramic replication techniques described below, the ceramic material must be in a moldable state, i.e. · it must fill the mold completely; · it must be consolidated in the mold; · it must be demoldable after filling so that all micorstructured details are kept and not destroyed; · it must be sinterable to the desired density without cracking or fragmentation. These requirements can best be met by colloidal techniques [1–3]. The following section deals with embossing of ceramic tapes, electrophoretic deposition, gel casting, direct coagulation casting and soft molding as promising techniques for the fabrication of micropatterned ceramic parts.

15.2 Embossing of Ceramic Tapes

Embossing techniques have been applied for the micropatterning of ceramic thin films down to the sub-micrometer range [3–5]. In this technique, mostly elastomeric stamps made via soft lithography are used to pattern liquid thin films derived from sol–gels or colloidal slurries. In this context, however, we will focus on ceramic microcomponents that are replicated by the embossing of ceramic tapes. The casting of ceramic tapes by the doctor-blade process [6] is widely used in electronics for the fabrication of thin substrates, multilayered ceramic packages and multilayered capacitors [7, 8]. In this process, the ceramic powder is dispersed in a solvent, mixed with suitable organic additives, e.g. binders, plasticizers and dispersants, and then cast to thin, flexible tapes. After evaporation of the solvent, the tapes are usually cut or punched to the desired shape, possibly laminated to form a stack of thin layers and subsequently sintered. Most slurries for tape casting are based on organic solvents, but also aqueous systems have increasingly been used [9]. The selection of additives plays an important role for the quality of the ceramic tape [8, 10–12]. Whereas binders provide the strength of the green tape, the role of the plasticizers is to impart sufficient flexibility to the tape and dispersants are added to stabilize the slurry and to control its viscosity.

15.2 Embossing of Ceramic Tapes

A combination of tape casting and embossing has been used for the micropatterning of ceramic components [13, 14]. Zirconia/alumina, silicon nitride, hydroxyapatite (HAP) and lead zirconate titanate (PZT) powders, dispersed in ethanol-based slurries, were used to cast tapes with a thickness of about 500 lm. Ceramic loadings of 80–90 wt.% (40–53 vol.%) were realized in dried tapes made of zirconia/alumina and PZT, respectively, while the ceramic loading was only 50–60 wt.% (25–30 vol.%) in the case of HAP and Si3N4. Up to four layers of tapes were laminated and subsequently embossed with microfabricated metallic dies or PMMA molds made by the LIGA process. These dies featured different three-dimensional patterns with lateral dimensions between 10 and 120 lm and a structure height up to 600 lm. To achieve micropatterning of tapes by embossing, the flexibility and flow behavior of the tapes had to be enhanced compared with conventional, commercially available tapes. This is normally done by adjusting the ceramic loading and a suitable addition of dispersants and plasticizers. For the desired high ceramic loadings and ample structure heights, the applied pressures were relatively high. By variation of the pressure in the range 6–20 MPa, a good filling of the mold could be obtained, resulting in micropatterns with an aspect ratio of up to 4. To facilitate the demolding after embossing, water or stearic acid was used as a mold release agent. While the metallic dies could be separated from the tape after embossing, PMMA molds with an aspect ratio of > 4 were used as a lost mold, i.e. they were pyrolyzed during debinding. Owing to the relatively low ceramic loadings, the linear shrinkage was in the range 20–25 and 30–35% for zirconia/alumina and Si3N4 parts, respectively. The shrinkage was inhomogeneous with slightly higher values in the z-direction, i.e. in the structure heights. This behavior, however, was also observed during sintering of conventional ceramic tapes [15]. In Fig. 15-1 a–c, some examples of zirconia/alumina and silicone nitride micropatterns are shown. For the molding of the microstructures in Fig. 15-1 b and c, the same micromilled silicon wafer was used as an embossing tool. In the case of zirconia, the grooves are replicated in the ceramic, whereas in the case of sili-

a

b

Fig. 15-1 Ceramic micropattern made by tape casting and embossing. (a) Zirconia/alumina microstructure after sintering, molded with a 300 lm high metallic die, consisting of truncated pyramids. (b) Sintered zirconia/alumina and (c) Si3N4 microstructure molded with a micromilled silicon wafer. Reprinted with permission from [14]

c

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con nitride these details are lost during sintering owing to the transformation of the a-Si3N4 to the b-modification. Similar experiments were performed with zirconia green tapes that were embossed with a polyimide wafer, featuring lines and trenches in the range 4– 24 lm, at pressures between 3 and 51 MPa [16]. Although no complete filling of the mold was achieved with dry tapes, trenches in the range 10–24 lm could be filled with tapes that were softened by partial or total saturation of the tape with MEK (methyl ethyl ketone). Because of the reduced shear strength of the system, the pressure necessary to fill the trenches could be reduced with increasing saturation of the tapes. Tape embossing has also been used for the fabrication of PZT ultrasonic transducers with 1–3 connectivity (one transducing direction of ultrasound in a three-dimensional body) and for the fabrication of a ceramic heat exchanger [14, 17]. For the molding of PZT rods with a side length of 150 lm and a height of 400 lm, the embossing pressure had to be increased considerably to achieve complete filling of the mold. Simultaneously, however, the demolding was aggravated, so that only 90% of the rods could be demolded. Only after a draft angle was introduced into the molding die could PZT arrays having 4900 square rods with a side length of 120 lm in the sintered state be realized (Fig. 15-2 a). The arrays were embedded in polymer and, after cutting, polishing and electroding, the composites were poled. Owing to a slight distortion of the arrays during sintering, the final height of the composites was reduced to about 360 lm. These distortions during sintering may be caused by an inhomogeneous green density induced by the embossing process. In Fig. 15-2 b, a catalytic insert for a ceramic microreactor is shown, which was fabricated by embossing of laminated lithium aluminate tapes [18, 19]. Affected by the embossing process, in the upper part of the channel walls a greater amount of porosity was observed than in the lower part of the walls or in the bottom plate.

a

b

Fig. 15-2 Sintered ceramic components micromolded by embossing of tapes. (a) PZT array of columns with a side length of 120 lm and a height of about 400 lm. (b) Lithium aluminate channel structure. (Unpublished)

15.2 Embossing of Ceramic Tapes Fig. 15-3 Microheat exchanger plate fabricated by embossing of alumina tapes. Reprinted with permission from [21]. Copyright Springer-Verlag 1998

In a similar approach, stainless-steel embossing tools were used for the molding of alumina green tapes to fabricate heat exchanger plates and gear wheels (Fig. 15-3) [20, 21]. To achieve a better plastic flow, the ceramic loading was reduced and the binder and plasticizer contents were increased compared with commercial tapes. As most binders exhibit thermoplastic behavior, the embossing temperature can also be used as a further parameter to optimize the mold filling. Walls having a width of 30 lm and an aspect ratio of 3 were successfully replicated, but the parts were slightly bent after sintering and a grinding process was necessary to obtain a proper form fit. Su and co-workers [22, 23] used an embossing method in combination with a lost polymer mold technique for the fabrication of 3D ceramic microstructures. In their approach a ceramic dough is prepared based on viscous polymer processing (VPP) [24, 25] and calandered to a tape with a thickness of about 1– mm, which is embossed into PMMA [poly(methyl methacrylate)] molds at a pressure of 10 MPa. After a drying period, the ceramic green body is immersed in a solvent to remove the PMMA mold. Different PZT microrod arrays with aspect ratios of up to 10 were fabricated by this technique [23, 26]. Fig. 15-4 shows two examples of PZT arrays with feature sizes down to 30 lm. The reported

a

b

Fig. 15-4 (a) Hexagonal PZT array from CNC micromachined PMMA molds. (b) PZT microrod array with an aspect ratio of ca. 3, replicated from an x-ray lithography PMMA mold. Reprinted with permission from [23]. Copyright Springer-Verlag 2002

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piezoelectric properties of the 1–3 composite were predominantly within the range of expectations [26]. The embossing of tapes is a cost-efficient shaping method for micromolding of ceramics. As long as mechanical demolding is necessary or desirable, it is suitable for fairly simple features with aspect ratios up to *5. In the case of higher aspect ratios or more complex structure details, the mold has to be removed by dissolution or pyrolysis.

15.3 Electrophoretic Deposition 15.3.1 Principle

In microsystem and microchemical engineering technology, the fabrication of ceramic components with microstructured surfaces is – owing to their physical and chemical properties – of particular importance. For excellent surface roughness, small structural details and high dimensional accuracy, the use of submicron and nanosized starting powders is necessary. This requires a manufacturing process which allows the handling of small-sized powders. The phenomenon of electrophoresis (EP) is the movement of charged particles through a liquid under the influence of an external electric field applied across the suspension [27]. The particles move in a direction opposite or parallel to the external current, depending on their charge, and are deposited on either the cathode or anode. EP is therefore limited to conductive substrates that can serve as an electrode, but complicated shapes are easily accommodated [28]. Electrophoretic deposition (EPD) is a colloidal shaping process that allows the processing of powders by using aqueous or alcoholic suspensions. The process of electrophoretic deposition involves two steps: the electrophoresis, i.e. the migration of charged particles in an electrostatically or electrosterically stabilized suspension by the effect of an electric field, and the deposition of the particles on the oppositely charged electrode [29]. EPD provides some important advantages such as short formation times, little restriction in the shape of substrates, a simple deposition apparatus [30] and low solids content of the suspensions. Furthermore, EPD produces very homogeneous parts with high particle packing density and low surface roughness. In comparison with conventional processing techniques, there are other advantages. Unlike in filtration processes, there is no movement of the solvent through the deposited layer. Therefore, high deposition rates even for particles in the nanometer range and thick layers can be achieved. Moreover, in contrast to slip casting, the deposition rate is not highly dependent on particle size [31]. The process of electrophoretic deposition was used first by Harsany in 1927 to deposit ThO2 and tungsten on a platinum cathode [32]. In 1969, EPD was used for the first time to produce industrially formed components of alumina

15.3 Electrophoretic Deposition Fig. 15-5 Principle of electrophoretic deposition (EPD)

ceramics. Since then, a multitude of systems such as metal salts, oxides, inorganic and organic paints, rubber, glasses and dielectric materials have been investigated in aqueous and organic media [33, 34]. Later, the fabrication of porcelain crucibles and cups by this method was reported [35]. Nowadays, the EPD process is a technique used worldwide for the processing of particulate layers such as composite layers, superconducting films, laminates, filaments and organic and bioceramic materials [36, 37]. For the direct generation of microstructured parts or of microparts by electrophoretic deposition, it is convenient to have the deposition electrode inversely micropatterned. A sketch of the principle of EPD into a patterned electrode is shown in Fig. 15-5. 15.3.2 Suspension Stability

A particulate surface charge is necessary for a stable suspension that can be used for EPD. The generation of these charges can be obtained by the following processes [38, 39]: 1. Electrostatic stabilization. The basic mechanism is the generation of charges on the surface of the suspended particles by dissociation or ionization of surface groups. In aqueous systems, H+ and OH– are the potential-determining ions. Hence the kind of charge depends on the pH of the suspension. Electrostatic stabilization is the principal stabilization mechanism in aqueous systems. In ethanolic systems the electrostatic stabilization of, for example, Al2O3, ZrO2 and La2O3 includes three steps: (i) adsorption of undissociated ethanol molecules on the basic surface sites of the oxide particles, (ii) dissociation of the adsorbed ethanol molecules by hydrogen ions transferring to the basic surface sites and (iii) desorption of C2H5O– anions into the solution, leaving the particles positively charged. For SiO2, protons transfer from the acidic surface sites to the ethanolic medium in the second step, resulting in negatively charged particles.

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2. Electrosteric stabilization. The basic mechanism is the adsorption of ionized surfactants. Carbon black, for example, becomes negatively charged by adsorption of anionic surfactants. These charged particles attract oppositely charged ions around the particle and form the so-called double layer. When the particles move, some of the ions within the double layer ‘shear off’ and the resultant potential between the surrounding medium and the shear layer boundary is called the zeta potential. Particles repel each other and the higher the zeta potential the more stable are the particles against coagulation. 15.3.3 Kinetics of Electrophoretic Deposition

The thickness of the deposited layer depends on the rate at which a deposit forms during EPD. In 1940, Hamaker observed that the deposited weight is linearly proportional to the charge passed [40]. He also observed that the voltage drop within the deposited layer can be very high. Therefore, he proposed that the amount of the deposited mass m(t) is proportional to the particle concentration of the suspension c(t), the surface area of the deposit A, the electrical field strength E(t), the deposition time t and the electrophoretic mobility le: Zt m…t† ˆ

fAle c…t† E…t† dt

…1†

0

Regarding deposition kinetics, a second correlation was proposed by Henry [41, 42], which relates the electrophoretic mobility with the zeta potential and some properties of the dispersing medium: lˆ

2er e0 ff …jrp † 3g

…2†

where er is the relative permittivity, e0 the permittivity of vacuum (8.85 ´ 10–12 A s/V m), f the zeta potential (V), j the inverse of the Debye length, rp the particle radius (m), g the viscosity (Pa s) and f(jr) the Henry factor. For large particles, the Henry factor becomes f(jr) = 1 and Equation (2) transforms to the Debye–Hückel equation [43]: lˆ

2ef 3g

…3†

For small, spherical particles the Henry factor becomes f(jr) = 1.5 and Equation (2) transforms to the Helmholtz–Smoluchowski equation:

15.3 Electrophoretic Deposition

lˆ

ef g

…4†

Constant-voltage EPD means that the potential between the electrodes is maintained constant. Under constant-voltage conditions, the current decreases with time because the deposited film continually increases the cell resistance. Consequently, the deposition rate decreases with time as the potential loss across the film increases. Constant-current EPD supplies a constant voltage drop across the suspension, so all particles arrive with the same velocity over the deposition time, i.e. the deposition rate is constant. Hence constant-current EPD avoids the limited deposition problem of constant-voltage EPD [38]. 15.3.4 Fabrication of Microstructured Components by Electrophoretic Deposition

The fabrication of microstructured ceramic components can be performed by electrophoretic deposition from stable aqueous or organic solvent suspensions as well as from precursor sols. The main problem associated with the use of aqueous suspensions for EPD is the electrolysis-induced decomposition of water above a DC voltage of about 1.4 V. The electrolysis results in a movement of ions towards the oppositely charged electrode, where recombination of the ions results in the generation of gases. If the deposition takes place at one of the electrodes, these gases can be incorporated into the deposit, resulting in large and irregularly distributed faults or bubbles. There are several possibilities to prevent the microstructures from incorporating bubbles. By the membrane method, developed by Clasen and co-workers [44–47], the electrophoretic deposition is carried out at a porous, ion-permeable unstructured or microstructured membrane which is located between the electrodes (Fig. 156). Deposition occurs at the membrane whereas the ions can pass through the pores of the membrane so that recombination of the ions and generation of gas bubbles occur at the electrodes and not at the deposited structure. Alternatively,

Fig. 15-6 Principle of electrophoretic deposition at a membrane

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Fig. 15-7 Microstructured electroconductive paraffin electrodes

the applied voltage can be kept lower than the decomposition voltage of the suspension medium. A convenient way to fabricate ceramic microcomponents by EPD is the application of electroconductive lost molds as deposition electrodes [48]. These electrodes, made e.g. from paraffin wax enriched with carbon black to make them electroconductive, can be generated by a two-step replication process from a master mold. In the first step the master mold – made by any microstructuring process – is copied into a silicone rubber model, which in the second step is then inversely replicated into the desired electrode structure. (For more information on this replication method, see also Chapter 12.3.) Different types of such microstructured lost paraffin wax molds are shown in Fig. 15-7. With respect to the mentioned properties and advantages of electrophoretic deposition, the most decisive factor is the packing density of the green bodies. A maximum particle packing density leads to minimized sintering shrinkage and minimized sintering warpage. According to the particle packing theory [49], the packing density is directly influenced by the particle size distribution of the starting powders. A maximum density is achieved when the large particles form a hexagonally packed lattice with the interstices completely filled by smaller particles. For bimodal size distribution, for example, the ideal ratio of coarse to fine

Fig. 15-8 SEM image of sintered triple mirror array of alumina ceramic

15.3 Electrophoretic Deposition

Fig. 15-9 Three-dimensional structure analysis by confocal laser scanning microscopy

particles is > 6.5. Thus, an optimized particle size distribution should provide a homogeneous green body with a packing density as high as possible. To verify this theory, analyses were carried out with ethanolic suspensions containing alumina powder mixtures with defined bimodal size distributions. The packing densities of the green bodies, achieved by EPD, were studied as a function of the fraction of fine powder. It could be shown that a maximum density of 77% can be achieved in the range 10–25% fine powder. A precise knowledge of the influence of particle size distribution on packing densities allows the fabrication of microstructured ceramics by EPD. Fig. 15-8 shows a sintered triple mirror array without visible cracks or warpage and a good molding accuracy. This good replication quality, especially with respect to the 908 angles of the isogonal surfaces, could be confirmed by confocal laser scanning microscopy as shown in Fig. 15-9. Microstructured substrates which are electroconductive can also be used for the EPD of silica microstructures [50].

Fig. 15-10 Uncoated substrate and coated substrate (left) and SEM image of sintered layer (right)

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Because of their chemical inertness, ceramics are also considered suitable as protective layers. Electroconductive surfaces such as microheat exchanger elements can be coated by electrophoretically depositing submicron and nanosized zirconia at low voltage (Fig. 15-10). The layers on microstructured steel substrates are dried in vacuum and sintered for 15 min at 1000 8C. For this, aqueous electrostatically stabilized suspensions containing nanosized zirconia particles are developed. The characterization and analysis of the surface and structure dimensions by SEM and white-light interferometry demonstrate that sintered ceramic coatings with a thickness of a few microns can be achieved. 15.3.5 Sol–Gel Electrophoretic Deposition of Microstructured Ceramics

A special process for the generation of three-dimensional microcomponents by EPD is sol–gel electrophoretic deposition [28]. By this process, piezoceramic microparts were produced directly from liquid metal–organic lead zirconate titanate (PZT) precursors [51–54]. The precursors were PZT sols, synthesized by a modified sol–gel technique. After EPD of these sols into a microstructured mold, simultaneously with the gelation of the sol, the deposited gel body was then dried and the mold was removed. The following step comprised the simultaneous transformation of the organics into the ceramic state by pyrolysis and the final sintering process. Special emphasis was laid on the development of the precursor [52, 53]. By introducing polymerizable chelating agents into the synthesis, the gel can additionally be hardened by radical polymerization. The use of UV-sensitive initiators allowed gel hardening by photoinduced polymerization, resulting in gels which showed considerably better mechanical strength than comparable inorganic gels owing to mutual interpenetrating inorganic and organic networks in the gel instead of only an inorganic network built up by gelation.

15.4 Gel Casting

Gel casting was invented and further developed at ORNL by Janney, Omatete and co-workers [55–62]. The principle is as follows. A concentrated slurry of ceramic powder in a solution of organic monomers is poured into a mold and polymerized in situ to form a green body in the shape of the mold cavity. Owing to the low viscosity of the monomer, the mold can easily be filled with the slurry and after polymerization the gel fixes the ceramic powder in the desired shape. After consolidation of the green body by gelation, the mold can be removed, the green body is dried (the solvent has to be evaporated), the organic binder is burnt out and the dense ceramic part is obtained after the final sintering process. A schematic overview of the processing steps is given in the flow chart in Fig. 15-11. Initially the gel casting process was developed using an or-

15.4 Gel Casting Fig. 15-11 Flow chart of the gel casting process (according to [62])

ganic route. Multifunctional acrylate monomers are dissolved in an organic solvent [55, 56]. After polymerization by free-radical initiators, these monomers form highly cross-linked polymer–solvent gels, which can be further processed as shown in Fig. 15-11. However, because of potential environmental problems with the removal and disposal of the organic solvent and because of the familiarity of ceramists with aqueous systems, a new water-based gel casting technology was developed [55, 57–59]. The search for water-soluble monomers led to the acrylamide gel system which is used in biotechnology for gel electrophoresis [55, 57, 58]. For gel casting, two types of organic monomers are required. The main monomer has a single double bond and, when polymerized alone, forms a linear polymer. The second monomer is the cross-linking monomer with at least two double bonds. A cross-linked polymer–solvent gel is generated when these monomers polymerize simultaneously. In the acrylamide system the monomer is acrylamide (AM) and the cross-linking polymer is N,N-methylenebisacrylamide (MBAM). The standard premix solution (before adding the powder) contains 15 wt.% monomers in water with a mass ratio of AM to MBAM of 24 : 1. Typical solid contents for gel casting with this system are at about 50 vol.%. The residual polymer content after drying is then about 4 wt.% [59]. The disadvan-

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tage of this system is the neurotoxicity of acrylamide. This is why Omatete and co-workers [59, 60] replaced the acrylamide with the less hazardous methacrylamide (MAM), again with MBAM as a cross linker. Various developments for gel casting ceramics with nontoxic or environmentally friendly polymers or biopolymers as gel formers have been performed. They can be categorized as cross linking of poly(vinyl alcohol) polymers by metal ion complexation [63] or by cross linking of biopolymers, e.g. chitosan [64], gelatin [65, 66], agarose [67] or proteins [68]. Many new gel formers, especially from biopolymers and low-toxicity monomers, are being still studied [69]. Gel casting has been shown to be a very versatile process for generating solid green bodies of complex shape and with fine details from a colloidal precursor. Not only can various ceramic materials be formed by this process, but it is also applicable to metal powder casting [60]. The main objective of most of the studies performed is the development or the optimization of a gel former in combination with the ceramic material under consideration. Other developments are performed in order to improve the drying process and thus to avoid surface spattering, which is a crucial processing step in the gel casting process [70, 71]. According to the literature, the most investigated material with respect to gel casting seems to be alumina (Al2O3) as it is the most important engineering ceramic. The first basic studies were performed at ORNL [55, 56, 58, 59, 61] (see Fig. 15-12) as part of the development of the gel casting technology. Other researchers have developed and tested new gel formers. Prabhakaran and Pavithran [72] used acrylic acid in aqueous concentrated alumina slurry for gel casting of alumina. They obtained strong gels and green bodies by gelling at 80 8C and by using ammonium persulfate as free-radical initiator for polymerization. Jia et al. [73] replaced acrylamide with sodium alginate, added calcium salt and a chelator to stabilize the suspension and started the gelation by adding hexadionic acid. By this method they obtained complex-shaped green bodies with a

Fig. 15-12 Gel cast Al2O3 gear wheels. Reprinted with permission from [62]. Copyright ORNL

15.4 Gel Casting

homogeneous microstructure, showing no apparent agglomerates or pores. They could sinter the green parts without special binder removal treatment to densities of about 99% of the theoretical density. ZTA (zirconia toughened alumina) parts were successfully gel cast by Liu et al. [74], who measured a bending strength of 792 MPa. Silicon nitride (Si3N4) as another important engineering ceramic was also shaped by gel casting [60, 75, 76]. Janney et al. [60] developed a low-toxicity gel casting system using MAM–MBAM gel formers for shaping of Si3N4 parts. Huang et al. [75] also used aqueous processing with AM and MBAM gel formers. However, to improve processing in water and also to improve the resulting mechanical properties of the shaped Si3N4 parts they coated the Si3N4 powder particles with Al2O3 and Y2O3, which served not only as oxidic coatings on the Si3N4 powder but also as homogeneously distributed sintering aids. Thus, after gas pressure sintering they achieved a bending strength of 840 MPa with a Weibull modulus of 15.55. They stated that these values are much higher than those of cold isostatically pressed samples prepared from the original Si3N4 powder. To avoid the problems of hydrolysis in water, Stampfl et al. [76] used solvent-based gel casting for the shaping of different Si3N4 parts. They not only shaped small Si3N4 turbine wheels with details in the sub-millimeter range (see Fig. 15-13), but were also successful in the fabrication of Si3N4 microcomponents (Fig. 1514) by gel casting into a plasma-etched silicon mold. For gel casting of metallic parts of 17-4 PH stainless steel, they used aqueous gel casting based on the ORNL technique. Also silicon carbide (SiC) parts with complex shapes were successfully molded by this technique [77, 78].

Fig. 15-13 Green (a) and sintered (b) silicon nitride turbine by gel casting. Reprinted with permission from [76] Copyright Elsevier Science

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15 Further Ceramic Replication Techniques Fig. 15-14 Sintered test structure (a) and green turbine wheel (b) made out of gel cast silicon nitride. Reprinted with permission from [76]. Copyright Elsevier Science

Not only engineering ceramics can be successfully shaped by gel casting but also functional ceramics. This was shown, e.g., by Guo and co-workers [79, 80], who performed gel casting with lead zirconate titanate (PZT). They used the AM–MBAM combination as gel formers and found that, compared with die pressing, gel casting provided a more homogeneous microstructure and more homogeneous piezoelectric properties of the PZT samples. Their gel cast samples also showed slightly stronger piezoelectric effects than die pressed samples. Apart from the field of technical application, hydroxyapatite (HAP) as a bioceramic can be shaped by gel casting, as shown by Varma et al. [81].

15.5 Direct Coagulation Casting (DCC)

Direct coagulation casting (DCC), developed by Gauckler and co-workers [82– 84], relies on the physical gelation of colloidal particles, not on the formation of an organic network as in gel casting. In DCC, an initially electrostatically stabilized colloidal suspension is filled into the mold and then destabilized by enzyme- or pH-catalyzed reactions to produce H3O+ or OH– ions or solubilized salt. This is done either by changing the pH of the suspension from a pH re-

15.6 Soft Molding Fig. 15-15 Al2O3 and SiC parts cast via direct coagulation casting. Reprinted with permission from [83]. Copyright Elsevier Science

gime with high surface potential to the isoelectric point (IEP) of the powder or by increasing the ionic strength in order to compress the electrostatic double layer of the powder particles. Investigations were performed to establish the DCC process by defining suitable self-decomposing or enzyme-catalyzed reactions. The best results for alumina were obtained with an aqueous suspension, stabilized by HCl in the acidic state. As the component that destabilizes the suspension, urea was added in combination with urease as an enzyme that promotes the destabilizing reaction. By this or by similar processes, for which the destabilizing components were optimized with respect to the actual material, a wide variety of ceramic powders could be processed successfully [82–85]. Among these are complex alumina parts and also SiC and Si3N4 parts, ZnO varistors or composite parts from ZTA [82–84]. As an example, Fig. 15-15 shows some of these DCC parts (Al2O3 and SiC). Studies on DCC with Si3N4 were performed also by Jung et al. [86]. They reported a resulting green density of the cast samples of nearly 56% and a very homogeneous microstructure and porosity distribution.

15.6 Soft Molding 15.6.1 The Soft Molding Technique and Its Variants

Soft molding or soft lithography is a microreplication technique using elastomeric silicones as molds for micropattern transfer. The mold material mostly used is polydimethylsiloxane (PDMS). This technique was developed mainly by Whitesides and co-workers [e.g. 4, 87–92] and is also described in a review by Heule et al. [93]. Several variations of this technique exist and a schematic overview is given in Fig. 15-16. The general advantage of this technique is the elasti-

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city and the low surface energy of the elastomeric PDMS molds, which allow an easy release from the patterned structure. By soft molding, details down to the single-digit micrometer scale or even finer could be realized [90]. It was demonstrated by Xia et al. [94] that molding with elastomeric molds is possible with resolutions on the nanometer scale. According to Fig. 15-16, replica molding (REM) differs from other molding principles by using elastomeric rather than rigid molds. Low-viscosity feedstocks, suspensions or polymer precursors are filled into the mold, solidified,

Fig. 15-16 Schematic illustration of the soft molding variations: (a) replica molding (REM); (b) microtransfer molding (lTM); (c) micromolding in capillaries (MIMIC); (d) solventassisted micromolding (SAMIM) (according to [90])

15.6 Soft Molding

e.g. by curing, followed by peel-off from the mold. This is the same technique as used in hot molding or low-pressure injection molding (see Chapter 12). By this method, free-standing micropatterns are achieved. In microtransfer molding (lTM), the filled mold is placed or slightly pressed on to a support or substrate. By this method, micropatterns on a substrate can be realized. In micromolding in capillaries (MIMIC), the mold is placed with the patterned side on a substrate and a suspension or a liquid precursor is driven into the micropatterns by capillary forces. After drying or curing, the mold is removed and the micropatterns are fixed on the substrate. Free-standing microstructures can be achieved by removing the substrate, e.g. by dissolving a sacrificial layer on the substrate or by etching the substrate away. Solvent-assisted micromolding (SAMIM) can generate quasi-three-dimensional microstructures on the surface of a polymer. The liquid precursor or suspension contains a solvent, which partly dissolves the polymer substrate without affecting the PDMS mold. After evaporation of the solvent, the PDMS mold can be peeled off and the resulting microstructures are fixed on the substrate. Various research and development work for the generation of ceramic micropatterns has been performed using the soft molding techniques. 15.6.2 Soft Molding of Preceramic Polymers or Precursors

In their study of the MIMIC process, Kim et al. [88] formed silica microstructures on a support using a solution of polyethoxymethylsiloxane with low viscosity (0.8 cP) as a precursor. After filling the capillaries, the solvent was evaporated at about 100– 150 8C before removing the mold. By curing at temperatures above 400 8C, the organic precursor was converted to SiO2. Beh and Xia [95] reported MIMIC as a useful method for micropatterning oxide ceramics such as ZrO2 or SnO2 on Ag-coated Si substrates starting from liquid organic precursors. After removal of the Ag layer by selective etching and after ceramization at temperatures of about 460 8C, they obtained free-standing ceramic structures with details in the range 2–4 lm. Also using MIMIC, Seraij et al. [96] molded Pb(ZrTi)O3 (PZT) and strontium niobate (Sr2Nb2O7) microstructures. They started from sols, which filled the capillaries. After gelation and removal of the PDMS molds, the gels were ceramized. In spite of the significant sintering shrinkage by this method, crack-free and uniform microstructures could be obtained after sintering at 700 8C (PZT) and at 800 8C (Sr2Nb2O7). Starting from self-assembling sol–gel precursors, P. Yang et al. [91] produced patterned continuous mesoporous oxide structures in microchannels by MIMIC. H. Yang et al. [92], also starting with organic precursors, produced borosilicon carbonitride (SiBNC) microcomponents (hexagonal grids and gear wheels) with details down to about 20 lm by both MIMIC and lTM (see Figs. 15-17 and 1518). To prevent the oxidation of the precursor, the patterning was performed in a glove-box under argon. As a specialty they used a hot-plate (*120 8C) for patterning in order to keep the precursor in a molten liquid state and in the case of MIMIC they used vacuum-supported MIMIC as described by Jeon et al. [97] to enhance the flow

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Fig. 15-17 SEM images of hexagonal grids of SiBNC ceramics. The widths of the hexagonal walls are (A) and (B) *20 lm and (C) and (D) 70 lm. (E) Ceramic structure at a tilt of *158 to the image stage; (F) the same structure after exposure to air at 1050 8C for 2 h; (G) sample as in (F) at higher magnification and a tilt of 608 to the image stage. Reprinted with permission from [92]

into the channels. Performing lTM, they filled the molten precursor into the mold and pressed a silicon wafer with a pressure of *3 ´ 104 Pa on top of the mold. Subsequently the filled mold was heated to 200 8C to cure the microstructures and finally the mold was cooled. After removal of the mold, the polymer patterns were transferred into the ceramic state in a covered gas furnace at 1050 8C.

15.6 Soft Molding Fig. 15-18 Micrometer-sized gear of SiBNC, patterned by lTM. (A) Top view of the gear; (B) side view of the teeth; (C) side view of the hole. Reprinted with permission from [92]

15.6.3 Soft Molding of Ceramic Suspensions

Another way to pattern ceramic microstructures by soft molding is to start with low-viscositys ceramic suspensions and not with preceramic polymers or precursors. Here the ceramic is already in its final composition and no phase transformation occurs. Investigations using this approach have been performed, e.g., by Zhang and co-workers [22, 98, 99]. As starting media they used aqueous alumina slurries with a solid content of 50–55 vol.%. In order to increase the solid content further they made a centrifugal casting of these slurries into the PDMS molds. After centrifugation, the mold was peeled off from the green parts, which were subsequently sintered. This method was reported to be applicable for microstructures showing aspect ratios < 3. It was verified by the fabrication of several alumina microparts such as micropistons [98], microgears [22, 98] and microconnecting rods [98]. The optimization of alumina slurries with respect to viscosity resulted in a highly concentrated suspension with a solid content of 84 wt.% alumina [99]. This was the starting material to perform micro-

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patterning by embossing using a PDMS mold or by lTM. Zhang et al. [99] showed that this stable and easily processible suspension was suitable for imprinting and lTM with PDMS soft molds. They achieved complete and dense microstructures on alumina substrates with resolutions around 100 lm and smallest details of 30 lm. Schönholzer and Gauckler [100] also studied the achievable accuracy of soft molding ceramic suspensions by micropatterning of alumina. They cast water-based suspensions of different aluminas and also of nanoscale boehmite with an average particle size of 45 nm into PDMS molds with cavities of 2 lm diameter. They found that the quality of the replicated patterns depends strongly on the particle size of the ceramic powders. Thus an alumina powder with 0.2 lm particle size gave better results than a powder with 0.5 lm particle size. They also found that the area of perfect pattern replication is strongly dependent on the drying shrinkage. This is why the sample starting with boehmite powder, which showed the sharpest replicated details, had an undestroyed area of only 20% of the sample. Because of the high initial porosity and the inevitable high sintering shrinkage, these samples could only be sintered to a density of 96% of the theoretical density, whereas the other samples reached full density. Soft molding of ceramic suspensions has been shown to be a versatile method for producing micropatterned ceramic sensor elements [101–104]. Heule et al. [101] formed tin oxide microlines only 10 lm wide by the MIMIC method, starting with an aqueous suspension of SnO2 powder with 280 nm particle size. Further developments were made by Schönholzer et al. [105], who patterned flat substrates with photoresist, thereby generating a micropatterned mold. They referred to this fabrication scheme as photoresist casting (PRC). Here the substrate acts as the bottom of the mold, and the photoresist structures represent the walls. After filling the structure with the ceramic suspension and after drying, the photoresist is removed and the patterned ceramic is sintered. These studies were the basis for the development of a complete SnO2 gas sensor element [102], fabricated by combination of photolithography and MIMIC (see Fig. 15-19). The PRC technique was also used for the fabrication of a miniaturized enzyme reactor with porous microstruts [106]. A microfluidic channel, 2000 lm wide and 15 lm deep, was segmented by 9000 of these struts with dimensions of 15 ´ 200 lm in order to split the flow in the channel. PZT microstructured sensor elements were also fabricated using the soft molding technique. Rosquist and Johansson [103] made piezo elements with crosswise layered channels of 40 lm width and 100 lm height. This was performed by generating micropatterned tapes by soft molding followed by lamination of the tapes. Further studies showed that also small PZT pillars 10 lm wide and 50 lm high can be made by soft molding replication. PZT columns with cylindrical and square geometry were molded by Gebhardt et al. [104] (see Fig. 15-20). The square columns had diameters of 81 lm, interpillar spacings of 36 lm and heights of 309 lm. The cylindrical columns had diameters of 48 and 656 lm and heights of 309 or 325 lm and were arranged in either a hexagonal or an irregular pattern. After

15.6 Soft Molding

Fig. 15-19 (a) Top view of a gas sensor setup; (b) isometric drawing; (c) microthick film gas sensors. Reprinted with permission from [102]

sintering, the arrays were filled with epoxy and the base plate was machined away. The authors characterized the piezoelectric properties. It could be shown that hexagonal or irregular arrangements are favorable.

Fig. 15-20 SEM micrographs of PZT array structures with cylindrical rods in a hexagonal arrangement and rod diameters of 65 lm and heights of 309 lm (top) and in irregular arrangement with rod diameters of 48 and 65 lm diameters and heights of 325 lm (bottom) patterned by soft molding. Reprinted with permission from [104]. Copyright Elsevier Science

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15.7 Conclusion

In this chapter, an overview was presented of several interesting replication techniques that permit the fabrication of ceramic microstructures and microcomponents. They cover a wide range of application possibilities and show special advantages. They all have in common that tooling is fairly inexpensive compared with high pressure injection molding. By lost-mold EPD, gel casting, and DCC, real 3D microcomponents can be realized, whereas by soft molding and tape embossing, 2.5D microcomponents can be generated, roughly comparable to injection molding. Soft molding allows the fabrication of microstructures and microcomponents with details in the single-digit micrometer range, but usually this very versatile technique needs a substrate or a base plate on which the structures are built. If this substrate is silicon, on-chip technology and the building-up of layered functional structures (e.g. for gas sensors) is possible with this technique. With respect to the potential of the mentioned techniques for large-scale production, tape embossing and soft molding, especially lTM, show this potential, because basically they can revert to existing printing and embossing techniques. This is comparable to the state of the microinjection molding technique, which in principle can revert to existing machinery for precision engineering. However, as described in the preceding chapters, development work still needs to be done for ceramic microinjection molding (lCIM) to meet the requirements for microfabrication. Also, the casting techniques and EPD can be scaled up to an industrially relevant scale. For future attempts to realize still finer details on the sub-micrometer or nanometer scale, there seems to be great potential for those techniques using preceramic organic precursors, e.g. soft molding with preceramic polymers or precursors or sol–gel EPD.

15.8 References 1 J. A. Lewis, J. Am. Ceram. Soc. 2000, 118, 2 3 4

5

2341–2359. F. F. Lange, J. Am. Ceram. Soc. 1989, 72, 3–15. C. R. Martin, I. A. Aksay, J. Electroceram. 2004, 12, 53–68. C. Marzolin, S. P. Smith, M. Prentiss, G. M. Whitesides, Adv. Mater. 1998, 10, 571–574. C. A. Bulthaup, E. J. Wilhelm, B. N. Huber, B. A. Ridley, J. M. Jacobson, Appl. Phys. Lett. 2001, 79, 1525–1527.

6 J. C. Williams, in: Treatise on Materials

Science, Technology, Vol. 9, F. F. Y. Wang (ed.); New York: Academic Press, 1976, pp. 173–198. 7 R. E. Mistler, in: Ceramic Processing, R. A. Terpstra, P. P. A. C. Pex, A. H. de Vries (eds.); London: Chapman and Hall, 1995, Chapter 5, pp. 147–173. 8 H. Hellebrand, in: Materials Science, Technology, Vol. 17A, Processing of Ceramics, Part I, R. J. Brooks (ed.); Weinheim: VCH, 1996, pp. 189–265.

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ter. Sci. Forum 2003, 442, 27–36. R. Moreno, Am. Ceram. Soc. Bull. 1992, 71, 1521–1531. R. Moreno, Am. Ceram. Soc. Bull. 1992, 71, 1647–1657. J. Böhnlein-Mauss, W. Sigmund, G. Wegner, W. H. Meyer, F. Hessel, K. Seitz, A. Roosen, Adv. Mater. 1992, 4, 73–81. R. Knitter, C. Odemer, E. Günther, U. Maciejewski, German Patent No. DE-PS 43 10 068, 1994. R. Knitter, E. Günther, U. Maciejewski, C. Odemer, cfi/Ber. DKG 1994, 71, 549– 556. D. J. Shanefield, R. E. Mistler, Am. Ceram. Soc. Bull. 1974, 53, 416–420. M. L. Griffith, A. R. Barda, N. Taylor, J. W. Holloran, in: Ceramic Processing, Technology, Ceramic Transactions, Vol. 51, H. Hausner, G. L. Messing, S. Hirano (eds.); Westerville, OH: American Ceramic Society, 1995, pp. 321–325. R. Knitter, E. Günther, C. Odemer, U. Maciejewski, Microsyst. Technol. 1996, 2, 135–138. R. Knitter, D. Göhring, M. Bram, P. Mechnich, R. Broucek, in: 4th International Conference on Microreaction Technology, Topical Conference Proceedings, AIChE Spring Meeting, 5–9 March 2000, Atlanta, GA; 2000, pp. 455–460. R. Knitter, C. Odemer, in: Material- und Verfahrensentwicklung für mikrotechnische Hochleistungsbauteile, K.-H. Zum Gahr (ed.), Wissenschaftliche Berichte FZKA 6528; Karlsruhe: Forschungszentrum Karlsruhe, 2000, pp. 151–155. M. Stadel, H. Freimuth, V. Hessel, M. Lacher, Keram. Z. 1996, 48 1112– 1117. V. Hessel, W. Ehrfeld, H. Freimuth, V. Haverkamp, H. Löwe, Th. Richter, M. Stadel, A. Wolf, in: Microreaction Technology, Proc. First Int. Conf. on Microreaction Technology, W. Ehrfeld (ed.); Berlin: Springer, 1998, 147–155. B. Su, D. Zhang, T. W. Button, J. Mater. Sci. 2002, 37, 3123–3126. B. Su, T. W. Button, A. Schneider, L. Singleton, P. Prewett, Microsyst. Technol. 2002, 8, 359–362.

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J. Eur. Ceram. Soc. 1999, 19, 271–275. 66 L. J. Vandeperre, A. M. De Wilde, J. Luy-

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ten, J. Mater. Process. Technol. 2003, 135, 312–316. Z. Xie, J. Yang, Y. Chen, Y. Huang, Br. Ceram. Trans.1999, 98, 58–61. O. Lyckfeldt, J. Brandt, S. Lesca, J. Eur. Ceram. Soc. 2000, 20, 2551–2559. M. Potoczek, E. Zawadnak, Ceram. Int. 2004, 30, 793–799. J. Ma, Z, Xie, H. Miao, Y. Huang, Y. Cheng, W. Yang, J. Eur. Ceram. Soc. 2003, 23, 2273–2279. J. Ma, Z. Xie, H. Miao, L. Zhou, Y. Huang, J. Am. Ceram. Soc. 2003, 86, 266–272. K. Prabhakaran, C. Pavithran, J. Eur. Ceram. Soc. 2000, 20, 1115–1119. Y. Jia, Y. Kanno, Z. Xie, J. Eur. Ceram. Soc. 2002, 22, 1911–1916. X. Liu, Y. Huang, J. Yang, Ceram. Int. 2002, 28, 159–164. Y. Huang, L. Zhou, Q. Tang, Z.-P. Xie, J. Yang, J. Am. Ceram. Soc. 2001, 84, 701– 707. J. Stampfl, H. Liu, S. W. Nam, K. Sakamoto, H. Tsuru, S. Kang, A. G. Cooper, A. Nickel, F. B. Prinz, Mater. Sci., Eng. A 2002, 334, 187–192. Z.-Z. Yi, Z.-P. Xie, Y. Huang, J.-T. Ma, Y.-B. Cheng, Ceram. Int. 2002, 28, 369– 376. Z.-P. Xie, Y.-B. Cheng, Y. Huang, Mater. Sci. Eng. A 2003, 349, 20–28. D. Guo, K. Cai, Y. Huang, L. Li, Z. Gui, Mater. Res. Bull. 2003, 38, 807–816. D. Guo, K. Cai, L. Li, C. Nan, Z. Gui, Ceram. Int. 2003, 29, 403–406. H. Varma, S. P. Vijayan, S. S. Babu, J. Am. Ceram. Soc. 2002, 85, 493–495. F. Baader, T. Graule, L. J. Gauckler, Ind. Ceram. 1996, 16, 36–40. L. J. Gauckler, T. Graule, F. Baader, Mater. Chem. Phys. 1999, 61, 78–102. L. J. Gauckler, W. Si, Th. Graule, F. Baader, J. Will, Ind. Ceram. 2000, 20, 102– 109. W. Li, H. Zhang, Y. Jin, M. Gu, Ceram. Internat. 2004, 30, 411–416. Y. S. Jung, U. Paik, C. Pagnoux, Y.-G. Jung, Mater. Sci. Eng. A 2003, 342, 93– 100.

15.8 References 87 E. Kim, Y. Xia, G. M. Whitesides, Nature

1995, 376, 581–584. 88 E. Kim, Y. Xia, G. M. Whitesides, J. Am. Chem. Soc. 1996, 118, 5722–5731. 89 M. Trau, N. Yao, E. Kim, Y. Xia, G. M. Whitesides, Nature, 1997, 390, 674–676. 90 Y. Xia, G. M. Whitesides, Angew. Chem. Int. Ed. 1998, 37, 550–575. 91 P. Yang, A. H. Rizvi, B. Messer, B. F. Chmelka, G. M. Whitesides, G. D. Stucky, Adv. Mater. 2001, 13, 427–431. 92 H. Yang, P. Deschatelets, S. T. Brittain, G. M. Whitesides, Adv. Mater. 2001, 13, 54–58. 93 M. Heule, S. Vuillemin, L. J. Gauckler, Adv. Mater. 2003, 15, 1237–1245. 94 Y. Xia, J. J. McClelland, R. Gupta, D. Quin, X.-M. Zho, L. L. Sohn, R. J. Celotta, G. M. Whitesides, Adv. Mater. 1997, 9, 147–149. 95 W. S. Beh, Y. Xia, J. Mater. Res. 1999, 14, 3995–4003. 96 S. Seraij, Y. Wu, N. E. Jewell-Larsson, M. J. Forbes, S. J. Limmer, T. P. Chou, G. Cao, Adv. Mater. 2000, 12, 1421–1424.

97 N. L. Jeon, I. S. Choi, B. Xu, G. M.

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Whitesides, Adv. Mater. 1999,11, 946– 950. D. Zhang, B. Su, T. W. Button, Adv. Eng. Mater. 2003, 5, 924–927. D. Zhang, B. Su, T. W. Button, J. Eur. Ceram. Soc. 2004, 24, 231–237. U. P. Schönholzer, L. J. Gauckler, Adv. Mater. 1999, 11, 630–632. M. Heule, J. Schell, L. J. Gauckler, J. Am. Ceram. Soc. 2003, 3, 407–412. M. Heule, L. J. Gauckler, Adv. Mater. 2001, 13, 1790–1793. T. Rosquist, S. Johansson, Sens. Actuators A 2002, 97–98, 512–519. S. Gebhardt, A. Schönecker, R. Steinhausen, W. Seifert, H. Beige, J. Eur. Ceram. Soc. 2003, 23, 153–159. U. P. Schönholzer, R. Hummel, L. J. Gauckler, Adv. Mater. 2000, 12, 1261– 1263. M. Heule, K. Rezwan, L. Cavalli, L. J. Gauckler, Adv. Mater. 2003, 15, 1191– 1194.

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V Automation and Quality Assurance

Advanced Micro and Nanosystems Vol. 4. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31493-8

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16 Automation of the Powder Injection Molding Process H. Weule, J. Fleischer, C. Buchholz, Institute of Production Science (wbk), University of Karlsruhe (TH), Germany Abstract Powder injection molding (PIM) offers great potential for the fabrication of micromechanical parts manufactured in metal or ceramic material providing a large variety of properties. To ensure economical microPIM production in large lot sizes and high-quality automation of the process beginning with demolding, handling, debinding and ending with sintering is a necessity. Within the field of automation the research focus is to optimize critical processes such as separation, demolding and handling and also the setting up of an autonomous and automated process chain, which are discussed in this chapter.

Keywords manufacturing; automation; handling; miniaturization; micromachining

16.1 16.2 16.2.1 16.2.2 16.3 16.3.1 16.3.1.1 16.3.1.2 16.3.1.3 16.3.1.4 16.3.2 16.3.2.1 16.3.2.2 16.3.3 16.3.3.1 16.3.3.2

Introduction 452 Microproduction Technology for Micromechanical Parts 453 Powder Injection Molding Process 453

Challenges of Automation for PIM in Microsystem Technology State of the Art of Automation Processes Handling of Fragile Microparts 455 Mechanical Grippers 456 Vacuum Grippers 457 Adhesion Grippers 457 Electrostatic Grippers 458 Separation of MicroPIM Parts 458 Tooling-internal Separation 458 Tooling-external Separation 460 Demolding of Microparts 461

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Mechanical Demolding Within the Tooling 461 Demolding by Nonmechanical Systems 461

Advanced Micro and Nanosystems Vol. 4. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31493-8

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16.3.4 16.4 16.4.1 16.4.2 16.4.3 16.4.4 16.4.5 16.5 16.6 16.7

Process Chain for MicroPIM Processes 462 Approach

463

Separation 463 Demolding 464 Handling 465 Automated Material Flow 467 Debinding and Sintering 469 Conclusion 471 Acknowledgments References 472

471

16.1 Introduction

Miniaturization is one of the most important technological trends, which will gain further importance in the future. It offers the decisive advantage of an increased function density on the smallest space and hence the possibility to develop new product concepts based on high integration of functionality. This opens up a wide range of new application fields associated with handy, portable products [1]. Two main approaches can be identified within the area of miniaturized products. On the one hand, there are silicon-based technologies and also lithography and electroplating processes which work on a 2.5-dimensional basis and originate in the area of microelectronics. On the other hand, there is the field of precision engineering, which is being subjected to increasing pressure to downscale the traditional production technology processes and thus qualify them for manufacturing miniaturized systems. The production of these systems is characterized by the assembly and combination of previously manufactured components [2]. In comparison with silicon-based products, a broad variety of materials are available which in turn provide a vast variety of property profiles (e.g. high tensile stress). As a consequence, they can assume mechanical functions, e.g. in microgears. Miniaturized systems with highly integrated functionality are required in different quantities – from mass production for, e.g., the automotive sector or medical applications to medium- and small-sized series in e.g. mechanical engineering. To ensure the economic production of microparts in large series, automation of the process has to be developed and integrated into an overall process chain.

16.2 Microproduction Technology for Micromechanical Parts

16.2 Microproduction Technology for Micromechanical Parts

One of the production technologies to meet the increasing industrial requirements of the economical production of high-quality parts in small to large series is the powder injection molding (PIM) process. It offers the production of three-dimensional microparts and also microstructured parts in a wide variety of metallic and ceramic materials. Owing to the strength of wear-resistant materials, highly precise metallic/ceramic parts for the transmission of forces and torques can be produced. As a result, PIM technology permits the economic production of microstructured parts and systems in various branches of industry such as the biotechnology, telecommunication and automotive sectors, etc., with throughputs of more than 1 million parts per year. However, in order to utilize this potential and to transfer this technology to industry, the complete automation of the microPIM process chain from the separation of the microPIM parts through sintering to stacking is essential [3]. 16.2.1 Powder Injection Molding Process

For the processing of metal powder or ceramic powder (diameter of powder d50 < 4 lm), it is first necessary to homogenize and granulate the material with a binder in a mixing unit. This compound, which is capable of being molded, is then injected into the cavity (see Fig. 16-1). After injection, the parts (green compact) have to be separated from the sprues, demolded and transfered to a furnace. In the subsequent debinding process, the binder is first removed from the molded part (green compact), resulting in the ‘brown compact,’ which is then sintered (for extensive descriptions of the microPIM process, see Chapters 10–12). Within these process steps, the reproducible and material-specific shrinkage of the part is typically 18–30% [4, 5]. 16.2.2 Challenges of Automation for PIM in Microsystem Technology

For several decades, the PIM process has been successfully used in macromolding for high-volume products and three-dimensional structures, e.g. toothedwheel turbine blades. Miniaturization of the parts produced leads to different challenging fields in the automation of the process chain [6]: · separation of sprues and parts within a tolerance below 1 lm and good surface qualities; · demolding of three-dimensional parts with structures below 50 lm; · batch handling of the fragile microparts under environmental conditions of e.g., high temperature and contamination of tooling and handling system by the feedstock;

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16 Automation of the Powder Injection Molding Process PIM-Process

Fields of automation

Machine Equipment for PIM

Separation

Demolding

Handling

Transport

Debinding/ Sintering Fig. 16-1 Visualization of the microPIM process. Source: wbk, OBE

· automation and linkage of the molding process and the thermal processes of debinding and sintering with minimal vibration; · continuous debinding and sintering with integrated quality assurance. To ensure both economic production and high-quality products, manual processes, e.g. handling, have to be replaced by automated processes. Critical success factors regarding automation are demolding, separation from sprues and handling of the parts. In contrast to macro dimensions, the maximum tensile stress acting on microstructures during the handling has to be within the range of millinewtons. The positioning precision has to be within the range 1–5 lm.

16.3 State of the Art of Automation Processes

In this section, the automation aspects of the microPIM process are described; for detailed descriptions of the state of the art of the microPIM process, see Chapters 10–12. In the field of microPIM, only a few research projects are known, apart from those at the Institut für Materialforschung IMF III of the Forschungszentrum Karlsruhe [7–9], which has activities in the fields of MIM, CIM and polymer injection molding.

16.3 State of the Art of Automation Processes

There are also research activities within the field of micro- and miniMIM processes at the Institut für Fertigungstechnik und Materialforschung (IFAM) of the Fraunhofer Gesellschaft in Bremen [10–12], the Fraunhofer Institut für keramische Technologien und Sinterwerkstoffe (IKTS) in Dresden and the Department of Engineering Science and Mechanics, Penn State University, USA [13]. Regarding the microCIM process, research work at the Institute für Mikrotechnik Mainz is well known, and in the field of micropolymer injection molding at the Institute für Mikrotechnik Mainz and the Institut für Kunststoffverarbeitung der RWTH Aachen [14]. Research activities for an automated production system for microPIM parts made of ceramic or metal have not been established so far. Within the field of plastic injection molding, different machine technologies especially for microdosages below 60 mg have been developed, but peripheral processes such as quality assurance and handling have not been the focus of machine development [15]. Micropolymer parts offer a higher resistance and stability regarding handling operations in comparison with microPIM parts in the green state. Therefore, handling results in conventional gripping technologies or extraction by vacuum. If oriented handling and stacking are required, microgrippers, which are described in the following sections, can be applied. Owing to the very low masses of the parts, quality assurance is a critical factor in an automated process chain. For statistical process control, a new optical tactile measurement system offers high potential. Owing to the restrictions of cycle times within the microPIM process, this offline measurement equipment cannot be integrated in the process chain for inline measurements. In addition to the challenges of contrast and focus, the optical measurement principle is difficult to apply (e.g. for critical functional structures, inner structures of three-dimensional parts with high aspect ratios), but nevertheless it is the most promising solution [16–18]. The state of the art of special fields of automation within the microPIM process chain is discussed in the following. 16.3.1 Handling of Fragile Microparts

For the handling of microPIM parts from the cavities to a stack, handling and gripping kinematics are necessary. Owing to the higher strength of polymers, the microplastic parts are demolded by use of ejector pins and transported by a regular conveyor belt. This technical solution does not find application within the automation of the PIM process owing to the low mechanical strength of the green compact. Industrial application of grippers such as magnetic grippers for MIM parts (5–10 mm smallest part length) or mechanical grippers for both MIM and CIM show considerable disadvantages. The main problem is the lack of force control of the gripping system, which is necessary for the handling of microPIM structures with maximum gripping forces in the millinewton range [19].

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Grippers for the assembly of the smallest parts can be classified according to four different effective principles: · mechanical grippers; · vacuum grippers; · adhesion grippers; · electrostatic grippers. This classification is also used for the description of the grippers that are under development or already commercially available in research and industry. A description is given in the following.

16.3.1.1 Mechanical Grippers A large variety of mechanical grippers for the assembly of the smallest parts can be found in the literature. All grippers consist of several joints and a drive. The design of the joints and drives is realized in very different ways. The plastic microgripper developed at the IWF, Braunschweig, has flexure joints and a shape memory actuator [20]. This gripper allows parts to be grasped with an edge length of up to 1 mm (see Fig. 16-2). In a further development of this gripper together with the IMT, Braunschweig, it was possible to minimize this gripper principle substantially by means of microtechnology. It is driven either by shape memory actuators [21] or by a pneumatic actuator that was newly developed for microgrippers [22]. The maximum size of the parts that can be handled with this gripper is 500 lm. The IPT, Aachen, has also developed a microgripper with flexure joints [23]. It is driven by two piezos. To enable the gripper to be adapted to any assembly task, it has a modular structure; the gripping jaws can be exchanged in a change station. The microgripper of the IWFT, Dresden, is also provided with flexure joints and is driven by shape memory actuators. The gripping forces of the different models are variable; the component length may amount to 3 mm. In the gripper for microassembly of the PAK, Kaiserslautern, the aluminum flexure joints of the gripper are pneumatically actuated [24]. This allows the gripping force to be varied according to the assembly task. The gripper produced by microtechnology of the IMMM, Ilmenau, is made of glass or silicon [25]. It is opened or closed by a piezo translator. The shape memory actuator developed at the FZK [26] and particularly the micropliers developed there to be used in surgery, which are driven by a piezo actuator, are produced by microtechnological processes. The same applies to two microgrippers of the BSAC, Berkeley, CA; one of them is actuated by a piezo and the other is opened and closed by an electrostatic actuator. A totally different approach for realizing an actuator is pursued by the IMES, Zürich. Their microgripper can be opened and closed by means of a thermal drive [27]. The microgripper of the LSR2, Lausanne, consists completely of shape memory material and uses the two-way effect for opening and closing of the gripper [28].

16.3 State of the Art of Automation Processes

The above-mentioned microgrippers were primarily developed in research institutes or institutions close to research. So far, grippers designed as prototypes have not been transfered into industrial applications. Grippers that are applied in industry are those which are used for larger parts, such as the pneumatically driven two-finger grippers by Schunk or Destaco. Depending on their structure, these grippers were in most respects adapted to the dimensions of the small components. Further mechanical grippers which are specially sold for use in the assembly of the smallest parts are frequently provided with flexure joints, but are mostly driven by a small electromotor (Bartels Mikrotechnik, Kleindiek Nanotechnik, Scientific Precision Instruments, Precision Instruments, Preiser Scientific, Emulation Technologie).

16.3.1.2 Vacuum Grippers Owing to the effective principle, the vacuum grippers used for the assembly of the smallest parts hardly differ from those used for larger parts. To allow the grasping of parts that are as small as possible, materials such as glass that have been rather unusual so far (SPI) or the latest technologies for the production of gripping nozzles are used (SPT Roth, Switzerland). Owing to the production technologies for small structures in wear-resistant materials for grippers, parts can be grasped which have a minimum side length of 10 lm (see Fig. 16-2).

16.3.1.3 Adhesion Grippers Adhesion grippers are grippers which are able to grasp and hold parts by means of an adhesive. Depending on the type of adhesive, three tendencies can be pointed out. First, the adhesive can be a liquid. For example the LS-MST, Cottbus, has developed a gripper in which a drop of liquid is brought on to a component which is gripped and held by means of capillary forces [29]. In this way, it is possible to grip successfully silicon chips of size 4 ´ 4 ´ 0.6 mm. Alternatively to the liquid form of the medium, it is possible to freeze the liquid as soon it is applied to the component (CSEM, Switzerland). The adhesive forces of the ice allow the holding of parts with up to 1 N/mm2 (this is a 50-fold higher holding force than that of the vacuum). For release, the ice is either broken or melted. The CRYOP by Naiss, an ice gripper working according to the same principle, has been developed in cooperation with the IWF, Berlin, and is already commercially available. The third alternative is to use a glue as adhesive that is applied on a strip. This has already been realized by the PAK [30]. After each gripping operation, the strip is moved further by a motor. The components are released by needles which push through the adhesive strip.

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16.3.1.4 Electrostatic Grippers Oh [31] has developed a gripper which is able to grasp and hold parts by means of an electrostatic field. Through the concentric set-up of an inner and outer electrode, it is possible to center the components (see Fig. 16-2). By an equally electrostatically driven actuator, the inner electrode can be moved so that the electrostatic field quickly breaks down, allowing immediate release of the parts. In a further development, it was possible to manufacture an electrostatic gripper adapted to the component sizes using microtechnological processes [32]. Additional alternative handling systems were developed in Munich to transport, lift and center contact-sensitive parts such as wafers via ultrasonic waves (acoustic levitation) [33]. A selection of the state-of-the-art grippers described above is shown in Fig. 16-2. In the field of high-precision handling systems, research activities at different universities show a trend towards assembly of fiber-optical and electronic systems [34, 35]. A parallel-tripod structure based on three coupled steel bands was built jointly by institutes in Munich and Braunschweig. The low-cost handling system offers high dynamics due to a low induced weight and an operational area 300 mm in diameter for, e.g., mounting of circuit boards [36]. At the ISR, Lausanne, and at AML, Minnesota, so-called l-factories were developed. The factories contain different machine tools which are interlinked by miniaturized handling systems with mechanical or vacuum grippers [19]. Both of these gripping and handling systems were developed especially for the assembly of micro(electronic) components in a small operating area. 16.3.2 Separation of MicroPIM Parts

Different challenges exist for the separation of microparts made by polymer injection molding and by microPIM. Owing to improved flow characteristics of polymers, a simpler design of the runner and the injection point is possible because injection points 40 lm in diameter and smaller are possible [15]. Injection points of microPIM parts have a diameter of 100 lm or more depending on the flow characteristics of the feedstock. There are different possibilities of separation: within the tooling via three-plate-tooling or hot-runner systems or outside the tooling via grippers or manufacturing technologies (ablation, milling or grinding). Owing to the main application of the possible solutions in industry, the following description is based on interviews and studies with experts.

16.3.2.1 Tooling-internal Separation The three-plate tool consists of different tooling plates – the runner system is within the first two plates and the cavities are within the second and third plates. By the use of a two-stage opening movement of the tool, sprue and parts are separated from each other and the gripping system can handle the parts from the ejection side of the tooling (see Fig. 16-3). Owing to the challenging

Fig. 16-2 Extract of the state-of-the-art of microgrippers

16.3 State of the Art of Automation Processes 459

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Fig. 16-3 Tooling-internal separation of parts and sprue (left, three-part tool; right, hot-runner system). Source: wbk, Kugele, Fa. m-Form

demands regarding the required precision of the tooling, this technology is well known in the field of macromolding but seldom finds application in micromolding [19]. The hot-runner system is based on a hot nozzle, which ends at the cavity of the tooling. The heated nozzle keeps the feedstock at melting temperature – therefore, the material can be injected at any time. As a result of challenges concerning thermal isolation of the nozzle and cavities and thermal damage to the feedstock owing to long periods at the melting temperature, this technology seldom finds application in the microPIM process [19].

16.3.2.2 Tooling-external Separation The most common technology for separating parts from sprue is tooling-external separation. Different solutions exist: separation via the gripping system, via cutting or ablation processes or via grinding. In the field of macromolding, the separation by gripper-integrated blades is successful in industrial application. After centering the gripper at the tooling, blades cut off the sprue from the parts while they are fixed in the tooling. For micromolding technologies this technical solution seldom finds application owing to potential damage to the cavities. Even cutting or laser ablation processes are used for macroparts but owing to the complexity and expense of positioning the parts outside the tooling for the separation process this alternative offers a technical but not an economical solution. Finally, separation may be done by grinding technologies. Here, microparts are molded on a substrate which can be easily demolded. After demolding and sinter-

16.3 State of the Art of Automation Processes

ing of the substrate, including the microstructures, the substrate can be removed by grinding. This time-intensive process involves high risks due to the small angular deflections between the grinding plane and the plane of the substrate [19]. 16.3.3 Demolding of Microparts

After cooling and solidifying, the microPIM part has to be demolded. Owing to undercuts, different shrinkages of part and tool, surface roughness of the cavities and adhesive forces, the part is held back in the cavity. The part shrinks material-specifically (1–3%) on inner cores or structures within the cavity, which increases the demolding forces. Owing to the main application of the possible solutions in industry, the following description is based on interviews and studies with experts. The technical alternatives for complete and nondestructive demolding depend on different parameters, e.g. geometry and size of the part, material, etc., and can be distinguished as follows.

16.3.3.1 Mechanical Demolding Within the Tooling Usually mechanical demolding systems are applied in industry and research. These systems consist of part-specific ejection pins and cores and are cylindrical or contour-specific as a part of the cavity. The actuation of this tool system is carried out by hydraulic or mechanical systems actuated by the injection molding machine by an enforced movement during the opening of the tooling. For special macroscopic parts, sliding systems are used in the case of undercuts or spindle units for threads. For macroscopic parts, a calculation basis and prototypical software systems exist to determine the demolding forces and thus to guarantee non-destructive and complete demolding. Furthermore, extensive experiments have been carried out to minimize the demolding forces by use of coatings or additional substances which are dosed before the molding process. Within the field of precision engineering and microsystem technology, these guidelines and calculations of demolding forces do not have complete validity and have to be adapted to conditions specific to microdimensions. Further, the characteristics of the two-phase-feedstock have to be taken into account. Research activities on the evaluation of demolding forces are not known [19].

16.3.3.2 Demolding by Nonmechanical Systems A special technology is demolding by pneumatic forces. This process is mostly used for macroscopic parts with large wall thicknesses and large production quantities. Here, pneumatic demolding can be used as a stand-alone technology or as a supporting mechanism for mechanical demolding. Within the field of precision engineering, this technology does not find application owing to possible damage to the fragile parts and a lack of orientation of the parts after demolding.

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Yet another technology is demolding by gripping systems using mechanical or pneumatic forces. To overcome the adhesion forces between the part and the cavity, demolding by gripping systems can only be applied if adequate gripping surfaces exist and therefore the parts protrude from the parting plane. Owing to missing surfaces or to possible angle failures, both of which can cause damage to the part, mechanical gripping systems are not applied for demolding. 16.3.4 Process Chain for MicroPIM Processes

To realize an automated process within the microPIM process, the following peripheral sub-processes have to be automated and incorporated into a process chain: stacking of green compacts, geometric quality assurance before and after the thermal processes, transport of the sinter trays in the debinding and sintering process and the integration of the thermal processes as continuous processes in the process chain. For stacking of PIM parts, industrial solutions exist for macro parts. Owing to their design regarding, e.g., shock resistance, gripping forces, etc., a transfer to microparts is not possible [37, 38]. In the microworld, the process chain consists of a manual change of magazines between the molding process and the thermal processes – automated solutions are not known. Even between the discontinuous batch debinding and sintering process, the change of sinter trays including the highly fragile brown parts is carried out manually. Fig. 16-4 shows the layout of an industrial production by injection molding machines with integrated handling systems. These handling systems find appli-

Fig. 16-4 State-of-the-art of a process chain in miniMIM. Source: OBE, Germany

16.4 Approach

cation for miniPIM parts larger than 3 mm. Automated handling of microPIM parts in medium to large series does not exist. After the molding process, the sinter trays are transported manually to separate batch furnaces for debinding and sintering owing to the lack of automation ability of batch furnaces. Even the part handling and the quality assurance after the sintering process are carried out manually [39]. In addition to the uneconomic handling with a deficit of discontinuous furnaces in comparison with continuous furnaces there is a lack of homogeneity and thermal stability. Research on decreasing the cycle time of continuous furnaces and the development of microcontinuous furnaces is not known [40].

16.4 Approach

The approach to the automation of the PIM process is divided into separation and demolding of microparts, handling and automation of fragile green compact parts produced by microPIM and integration of the debinding and sintering process in the complete process chain developed at the Institute Production Science (wbk), University of Karlsruhe. 16.4.1 Separation

Owing to the requirements regarding the surface quality and the precision of structure elements, the separation of the microparts, which are produced in a batch of four to eight parts, from the sprues is very critical. There are different possible ways of separation in the macroworld but they become challenging on entering the microworld. In principle, macroparts can be separated either by means of grippers in the tool or by using an external separation module. For microparts the external separation is complicated owing to the fragility of the connecting points of parts and sprues and the difficulty of positioning the parts in an external module within the range of 1 lm and below in order to meet the requirements of tolerances and surface quality. The alternative of separation within the gripping and handling module is difficult to adapt to microparts owing to possible tool and cavity damage. This leads to a defective molding and demolding process [41]. Therefore, attempts were made to realize a three-part tool, well known in the field of macroparts, for the manufacture of microPIM parts. In the three-part tool, the tool works in a two-stage movement. In the first stage, both sides of the cavities move from the fixed plate and thereby separate the parts from the sprues at the injection point. In the second stage, both plates with cavities open and the demolding and handling process can take place. A problem within this method is the small injection point, which may cause the feedstock to segregate into binder and powder.

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Fig. 16-5 Schematic view of a three-part tool for the separation of microPIM parts and sprues

The separation process has been realized using the example of a toothed wheel with an outer diameter of 1400 lm, a thickness of 230 lm and a module of 0.1 (involutes). The results show that the injection area realized has reached surface qualities meeting the design specifications and shows nondestructive feedstock characteristics with a centered gap of 0.04 mm2 [42]. The principal set-up of the three-part tool with a cavity of a toothed wheel within the millimeter range is shown in Fig. 16-5 (for better illustration, only a one-cavity tool is shown). Further developments will determine the minimal injection area without any segregation of the feedstock. Additional systems for separation within the tooling, e.g. ultrasonic, will be tested in the future. 16.4.2 Demolding

When the cavity is opened, the part has to be demolded from the tool. This process can be done with either mechanical ejector pins, application of pneumatic pressure within the tool or by the handling system. Owing to the microstructures, complete demolding is very critical. Adhesive forces between the feedstock and the cavity can exceed the material strength. Especially due to the shrinkage (about 2% of the nominal value) of the part during the cooling time, the part may stick on tool cores, for example. Research work at the wbk aims to evaluate demolding forces with respect to boundary conditions, e.g. necessary surface qualities, tribological characteristics, demolding slants of the mold to determine the position and size of ejector pins in the tool, as a basis for technical designers and toolmakers. For these objec-

16.4 Approach

Fig. 16-6 Demolding forces at microPIM test structures

tives to be achieved, correlations have to be set up between variable structures, sizes and alternative technical solutions (e.g. pneumatically) for demolding parts and structures by means of geometric sets. These sets include structures such as circles, triangles and hollow cylinders to model real parts, e.g. toothed wheels (see Fig. 16-6). The dimensions are below 300 lm with part lengths from 50 to 1000 lm and depths from 5 to 400 lm. Subsequent to this process, all parts are measured to detect damage. The experiments show small demolding forces below 1 N including an extensive reduction of the forces by use of a demolding spray and coating on the cavity for better surface qualities [43] but without further analysis of the effects of sprays on debinding and sintering processes. The described correlations will be completed by further experiments with variations in geometry. Within the automated process chain, demolding has been realized by mechanical ejector pins having an inner diameter of 525 lm, surrounding the inner core of the toothed wheel (see Fig. 16-5). 16.4.3 Handling

After demolding the parts from the cavities, the handling system has to transfer the parts to the stack unit. There are two critical process steps: · accurate positioning of the handling system to the cavities within a range of 1–3 lm in six degrees of freedom; · gripping of the fragile microPIM parts – a molded feedstock has a tensile strength of about 4 N/mm2, which restricts the gripper forces acting on the microparts to the millinewton range [37].

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Especially when extracting and gripping the parts from the cavity, high-precision positioning has to be ensured to prevent part damage. Compared with other high-precision applications (e.g. packaging and bonding processes in the electronics industry), the positioning of the handling system has to take place under environmental conditions of DT >> 40 8C, with possible contamination of the tooling and the handling system by use of feedstock/powder and vibrations from the injection molding machine. Further boundary conditions are phenomena that are specific to the microworld: during the handling, the gravitation force and the acceleration forces of the part are the determining forces. Both the gravitation force and the acceleration forces depend on the mass and hence on the volume of the component. Forces generated by pressure result in appropriate effective areas. During the size reduction, the surface forces decrease according to second order and the volume forces third order, so microparts with weights below 0.2 g and areas below 1 mm2 do not detach from the gripper owing to adhesion forces [44]. Based on these restrictions, different grippers for application in the microPIM process were investigated and developed at the wbk. Grippers based on the Venturi principle (pneumatic actuator and pneumatic handling of microPIM parts) and also magnetic grippers for metal injection molding (MIM) came into operation. The pneumatic gripper (see Fig. 16-7), based on the Venturi principle, consists of a housing with modular pestles for part variation. The movement of the pestle, essential for gripping the part at the cavity and releasing the part at the sinter tray, is realized by use of air bearings with slip-stick-free movement within the millimeter range. The gripping forces are controlled by air sensors. The mechanical gripper (see Fig. 16-7) is a two-finger gripper with two separated actuators which consists of flexure hinges for slip-stick-free movement and 5 mm motors. The gripping forces of the C167 microcontroller-based sys-

Smoory motor

Fig. 16-7 Schematic view of a pneumatic gripper with air bearings and a mechanical gripper with flexure hinges and force control

16.4 Approach

tem can be controlled by the use of strain gauges. The parts’ dimensions which can be handled with this system are between 100 lm and 3 mm. Substantial tests have shown reliable handling of the parts – in the case of magnetic grippers by the use of reverse polarity. Owing to adhesive forces between feedstock and tooling, which are higher than gripper forces, mechanical ejector pins were used in the demolding process. For handling the parts, the gripper is provided with an integrated lifting mechanism by means of which the part is extracted from the mold and released on the sinter tray. Additional gripping principles for higher and infinitely variable gripping forces will be developed, e.g. Peltier grippers (by freezing the parts to the gripper) and mechanical grippers (Fig. 16-7). Owing to process-specific boundary conditions, contactless handling systems, e.g. ultrasonic or electrostatic-based principles, cannot find application in the microPIM process [33]. 16.4.4 Automated Material Flow

To ensure reliable and complete demolding of the part, integrated quality assurance is necessary. Owing to the process restrictions, e.g. cycle time, a 100% visual inspection of the part is made after demolding and handling of the part to the sinter tray. Damaged or missing structures of below 1 lm can therefore be detected. Combined with the quality assurance, the parts are magazined on a sinter tray. Because so far sintering support devices have not been necessary for the current products and the position of the parts changes during the sintering process, a directional magazining with a precision of below 100 lm is implemented in the process. High-precision handling of the parts with a repeat accuracy of below 4 lm in six degrees of freedom is only needed for optical-tactile quality assurance and stacking after sintering. Owing to the sensitivity of green and brown compacts to vibrations, there are high machine requirements for conveyance systems. Furthermore, there has to be a flexible flow of material to ensure optimal sintering and charging of the furnace regarding different feedstock specifications and part/structure geometries and sizes. To meet the requirements, a paternoster and linear modules with a high gear reduction and a vibration absorber were installed in the system. A prototype of the process chain between the injection molding machine and the furnace is shown in Fig. 16-8. Stacking is done by means of a mechanical stage with a repeat accuracy of 50 lm – the movement in the third dimension is effected by the handling system. A linear module transfers the sinter trays to a paternoster for flexible storage. Up to three injection molding machines can be installed at the paternoster for furnace loading. A barcode system is integrated in the linear module for product identification. The sinter trays are clustered with the help of a production planning and control system and transferred to the furnace conveyer belt. After transferring the sinter trays to the furnace, they are piled up by means of a two-

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Paternoster for debinding process

Fig. 16-8 Schematic of the process chain for microPIM parts between molding machine and furnace

axis linear transfer module which handles the sinter trays and also spacing beams for stacking three sinter trays in one charge tray [45]. Subsequent to the debinding and sintering process, the sinter trays are extracted by means of a linear module with a hexapod mounted below. Therefore, the linear module transfers the sinter trays and the hexapod with a repeat accuracy below 1 lm handles the parts by use of a mechanical gripper developed by the Institut of Werkzeugmaschinen und Fertigungstechnik in Braunschweig [19]. Owing to the integrated vision system (see Fig. 16-9), the parts can be found, centered and handled by the mechanical gripper for stacking on 0.2 mm pins with a total system accuracy below 2 lm in all six axes. The scheme of the described process of stacking including the vision system and the centering process of gripper and part is shown in Fig. 16-9 (flow chart of vision inspection/ quality assurance of molded part, comparison with ideal part, centering of part within vision system and centering of gripper to part).

16.4 Approach

Fig. 16-9 Hexapod kinematics for stacking of microPIM parts (top) and quality assurance including centering process of gripper and sintered part (bottom)

16.4.5 Debinding and Sintering

To provide automated production of microPIM parts, the thermal processes of debinding and sintering have to be integrated. Generally, continuous and discontinuous furnaces are used for producing macro parts. In contrast to discontinuous (batch) furnaces especially for MIM, which are mostly used in laboratory environments, the most economic solution for large lot sizes is the use of a continuous furnace. Apart from their minor maximum capacity, batch furnaces have clear disadvantages in comparison with continuous facilities:

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· required loading of furnaces (debinding and sintering) sub-optimal for automation; · required transport of microPIM parts between debinding and sintering due to separate facilities. Especially the handling of debinded parts (brown parts with a binder fraction of < 3%) is a critical issue and not seldom a major cause of defects after sintering. As can be seen in Fig. 16-10, the dimensions of the furnace have a great influence on the consistency of quality. In a large batch-type furnace, great effort is required to guarantee homogeneous temperature and process conditions all over the furnace space. In a continuous furnace, the dimensions X and Y of a furnace are much smaller than Z and therefore the temperature or atmosphere differences along the width and height of the furnace can be neglected. Furthermore, the holding time and position Z in the furnace are fixed by the throughput rate. As a result, the number of process parameters that have to be controlled in a continuous furnace is only half of that in the batch furnace. Owing to the geometric proportion shown, a constant and homogeneous temperature within ± 3 8C is not achievable with a batch-type furnace. As a result, the part can be damaged and shape and positional tolerances may not be met. Therefore, in cooperation with an industrial partner, a continuous furnace was developed for debinding and sintering microMIM parts via catalytic or thermal debinding. A schematic of the realized but not yet tested furnace is shown in Fig. 16-11 [43]. The furnace (see Fig. 16-11) consists of a catalytic and thermal debinding area (maximum temperature: 600 8C), a sintering zone (maximum temperature: 1450 8C), a cooling area and a loading and unloading area. The total dimensions of the furnace are shown. The maximum capacity of the furnace (atmosphere: nitrogen and hydrogen in variable mixing ratios) is about 9 million parts/year with total cycle times below 9 h.

Fig. 16-10 Parameters with impact on the quality of thermal process and PIM parts

16.6 Acknowledgments

Fig. 16-11 Schematic of the continuous furnace for microMIM parts

16.5 Conclusion

The demand for complex three-dimensional micromechanical parts in wear-resistant material in different branches such medical, biological and telecommunication areas can be satisfied by the PIM process [38]. To obtain a breakthrough in this technology, economic and high-quality production of medium to large series has to be set up. This can only be achieved with an automated process and with special focus on critical sub-processes, e.g. separation, demolding and handling. To validate the concepts of automated microPIM production, a process chain for microPIM parts – here toothed wheels with a diameter of 1400 lm – has been set up. Innovative separation solutions for microPIM parts with a threepart tool and also demolding of microstructures, quality assurance, handling and transportation have been integrated in this autonomous process chain. To counter problems arising from producing smaller parts, dependences between structure sizes, material and specific technology criteria, e.g. gripping principles or force control systems, have to be developed as a basis for the respective subprocesses [46].

16.6 Acknowledgments

This research was financially supported by the German Research Foundation (DFG) within the Collaborative Research Centre (SFB) 499, ‘Design, Production and Quality Assurance of Primary Shaped Microcomponents Manufactured in Ceramic and Metallic Materials.’

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17 Microassembly – Approaches to Meet the Requirements of Accuracy J. Hesselbach, J. Wrege, A. Raatz, K. Heuer, S. Soetebier, Institute of Machine Tools and Production Technology (IWF), Technical University of Braunschweig, Germany Abstract This chapter presents an overview of the research activities carried out by the Institute of Machine tools and Production Technology (IWF) in the field of microassembly. These works are mainly based on the context of the collaborative research center 516 (SFB 516). General assembly concepts are discussed, performance specifications referring to accuracy are explained and different approaches to build up robots suited for microassembly are pointed out. Furthermore, three different parallel robots developed by the IWF are introduced and the possibility to improve the repeatability by integrating flexure hinges in the kinematic structure is discussed. Aspects of microgrippers and sensors which have to be integrated in an assembly system lead over to the description of two different assembly scenarios.

Keywords microassembly; assembly concepts; parallel robots; flexure hinges; microgripper; sensor guidance

17.1 17.2 17.3

Introduction 476 Assembly Concepts 477 Equipment Technology for Microassembly

17.3.1 17.3.2 17.3.2.1 17.3.2.2 17.3.2.3 17.3.3 17.3.3.1 17.3.3.2 17.3.3.3

Performance Specifications for Industrial Robots 478 Robots Suited for Microassembly 480 MICABOE Planar Parallel Robot [10] 481 TRIGLIDE Spatial Parallel Robot [11] 481 MICABOF Hybrid Parallel Robot [12] 483 Pseudo-elastic Flexure Hinges in Parallel Robots 483 Pseudo-elastic Flexure Hinges 485 Planar Compliant Robot (MICABOES) 486 Spatial Compliant Robot (TRIGLIDES) 487

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Advanced Micro and Nanosystems Vol. 4. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31493-8

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17.3.3.4 17.3.4 17.3.4.1 17.3.4.2 17.4 17.4.1 17.4.1.1 17.4.1.2 17.4.1.3 17.5 17.6

Experimental Results 489 Grippers for Microassembly 490 Mechanical Gripper 490 Electrostatic Gripper 491 Sensor Guidance in Microassembly

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Microassembly Scenarios 492 Assembly Scenario I (Sysmelec Autoplace 411) 492 Assembly Scenario II (MICABOF) 494 Comparison of the Microassembly Scenarios 497 Conclusion 497 References 498

17.1 Introduction

Worldwide, ongoing miniaturization of products and components in nearly all application areas can be observed. Not only are existing products being miniaturized but also a great number of new technical systems are being developed that would not be realizable without miniaturization. Products manufactured by using microsystem technology are predicted to have a continuous growing market potential. The NEXUS market analysis [1] anticipates a yearly growth of the world market for products based on microsystem technology of 20%. While the focus of the work within the SFB 499 is the production of single micromechanical components using advanced replication techniques, the SFB 516 ‘Design and Manufacturing of Active Micro Systems’ develops technologies for microsized electric linear actuators. These include design and simulation, manufacturing processes and bearings and sensors for active microsystems. A prerequisite for the integration of these different technologies is the availability of adequate assembly tools for the different product components. While some components are fabricated using photolithography, etching and electroplating techniques, others are made by conventional machining. In the final stage, the assembly of miniature-sized parts with an accuracy in the micrometer range is required. Hence adapted equipment, i.e. robots, grippers and feeders, is needed in addition to suitable strategies for the assembly process such as sensor guidance and quality assurance. The field of microassembly was not only opened up by technical advances in research projects but also by new industrial applications such as radar sensors for adaptive cruise control and acceleration sensors for airbag systems. In the beginning, the challenge was to increase the accuracy of pick-and-place machines for electronic production. Very small surface mount devices (SMD) and flip-chip components with dimensions of several hundred micrometers had to be handled. A major difference between microassembly and the pick-and-place process of SMD components in electronic production is the complexity and the required

17.2 Assembly Concepts

accuracy of the assembly process. Whereas in electronic production parts are placed on an adhesive surface in a 2D assembly with accuracies of 5–30 lm, target accuracies in microassembly are in the region of 1 lm. The complexity of the assembly process increases owing to the non-standardized range of components, the assembly of movable and 3D-shaped parts and the integration of adapted test equipment. These challenges have to be solved with consideration of economic aspects. To summarize, the essential microassembly problems are the development of suitable concepts, joining procedures and the necessary device technique. This chapter gives a survey of developments at the Institute of Machine Tools and Production Technology (IWF), Technical University Braunschweig, referring to microassembly.

17.2 Assembly Concepts

Potentials for reducing the assembly-involved costs are seen in production concepts that are adapted to the number of pieces produced in a product and in a component design that is in line with the assembly and, not least, with the development of equipment technology that is adapted to the requirements of microassembly [2, 3]. In contrast to the surface mounting processes of the packaging technology of electronics in which an automation adapted to the number of pieces produced is state of the art today, many problems regarding the production of miniaturized systems are not yet solved. Sensitive electrical, optical, mechanical and fluidic components with possible unfavorable geometries have to be stored in magazines, transported, picked and placed, grasped and processed. To take the variety of application-specific microsystems into account, strategies are currently being developed to standardize microsystem components and possible modular designs of microsystems [4, 5]. Two approaches are pursued as principal assembly strategies: parallel assembly using batch feeding and the serial assembly of discrete components (Fig. 17-1). In parallel assembly, many single objects are connected with each other at the same time directly without prior separating of the batch-manufactured objects. The resulting multi-layer compound component is then separated into individual microsystems. In parallel assembly strategies, attemts are made to transfer the batch manufacturing process from microtechnology to assembly in order to make use of its cost- and time-related advantages. These advantages arise from the fact that the state of order of the components can be maintained within the total process chain and that predominantly relatively large objects have to be handled. For this reason, some of the typical problems of discrete microassembly do not occur and it is possible to use conventional high-precision equipment technology. In most cases, however, serial assembly in microsystems production is inevitable since the high functional density can often only be ensured by combining

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Fig. 17-1 Schematic representation of different assembly concepts

the most different manufacturing processes for which a parallel assembly is excluded or only profitable if large numbers of parts are produced. At present, however, the serial assembly strategy still requires a relatively long assembly time. An alternative is a combination of serial and batch assembly concepts: serial batch assembly. With this type of assembly, it is possible to carry out process steps such as coating and cleaning or necessary subsequent steps such as housing in a batch process and thus in a cost-efficient way. Especially in the field of discrete assembly of microcomponents, an urgent need for research exists, focused on automated assembly. In this context, some solution approaches in the field of handling techniques are presented in this chapter.

17.3 Equipment Technology for Microassembly 17.3.1 Performance Specifications for Industrial Robots

The industrial robot is one of the basic devices to build up an automated assembly system. The suitability of robots for microassembly tasks depends on the performance specifications regarding accuracy. Usually different terms are used to describe these specifications. Slocum explains these terms in a demonstrative way [6]:

17.3 Equipment Technology for Microassembly

· accuracy is the ability to tell the truth; · repeatability is the ability to tell the same story each time; · resolution is the detail in which you tell a story. The ISO standard EN ISO 9283 [7] gives a definition for these terms and specifies appropriate test methods. To discuss accuracies with regard to the suitability of robots or the influence on microassembly processes, all terms should be used referring to EN ISO 9283. In the following they are explained using an abstract assembly task. Two parts have to be placed in an exact position relative to a substrate. For simplicity, it is assumed that the microparts are provided for a robot by the same feeding station. In order to provide the functionality of the assembled product, it is necessary for the parts to be in an exact position on the assembly surface and in an exact position relative to each other. The first action in the assumed assembly scenario is to move the endeffector of the robot to the pose of the feeding station in order to pick up the first part. The endeffector will not reach the desired pose exactly but a pose nearby which will be called actual pose. In subsequent approaches to the desired pose, the endeffector will reach a slightly different actual pose each time. The distance between the desired pose and the centroid of the subsequent actual poses is called absolute accuracy (AP). Another important value is the repeatability (RP), which is computed according to the equation RP ˆ I ‡ 3Si

…1†

whereI denotes the mean deviation from the centroid of subsequent actual poses and Si denotes the standard deviation. Graphically the repeatability can be described as the radius of a sphere in which 99.7% of all subsequent actual poses lie. The endeffector of the robot does not reach the desired pose exactly, nor does the fed part. The maximum error dmax, possibly occurring between the actual pose of the endeffector (E) and the actual pose of the fed part (F), is computed according to dmax ˆ …APE ‡ RPE † ‡ …APF ‡ RPF †

…2†

For a secure gripping of the micropart, the difference between the actual poses has to be compensated. This can be achieved either by a gripper with a self-centering effect and a corresponding gripping width or by sensor guidance. Moving the gripped part to its assembly pose involves similar problems to moving the endeffector to the pick-up pose. Again, the pose will not be reached exactly, but with an average distance AP between the desired and the actual pose. In the case of the assembly of parts providing a centering effect, a relatively high AP is tolerable. In other cases, even small deviations from the desired pose could lead to a malfunction of the assembly. In this case a sensor guidance of the assembly process is necessary in order to accomplish a stable process.

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If information from an adequate sensor is available, the current error can be fed back to the robot control and corrected afterwards. The quality and quantity of this correction depend on the resolution of the robot. The resolution is defined as ‘the smallest programmable distance, after whose demand a correct corresponding movement of the robot takes place’ [8]. If the current error between the pose of the endeffector and the desired pose were higher than the resolution, the error could be minimized to the quantity of the resolution. An error smaller than the resolution cannot be compensated. In cases where the endeffector has to move to the same position from different directions, e.g. if parts were to be assembled which are provided by two different feeders, the variance of the multiple direction position accuracy (vAP) would be of interest. This quantitiy expresses the deviation between the different average actual poses which result from approaching the same desired pose from three directions orthogonal to each other. Overall, it can be concluded that in microassembly the repeatability and the resolution are of the utmost importance. Depending on the application, the absolute accuracy and the variance of the multiple direction position accuracy can also be important. For benchmarking purposes and error analysis of robot kinematics, the exact defined methods and terms are essential. The stochastic migration of ball cages in ball bearings, for example, can lead to significant differences between the values of repeatability and the variance of the multiple direction position accuracy. 17.3.2 Robots Suited for Microassembly

A number of commercial robot manufacturers and many research institutions are developing robots which have a sufficient positioning accuracy for microassembly tasks [9]. In general, two different approaches are known: robots based on serial kinematic structures (serial robots) and robots based on parallel kinematic structures (parallel robots). Within these directions, again two approaches can be distinguished: the top-down approach and the bottom-up approach. In the former, researchers try to reach the required accuracies by modifying existing structures (e.g. adding redundant axes with high resolution), whereas in the latter one tries to build systems optimized for high precision from scratch. Serial robots are based on open kinematic chains, where all joints have a single degree of freedom each. Most serial robots for microassembly use the Cartesian structure. Often they are build using modular precision linear axes. In nearly all cases, direct measurement systems are used in order to rule out inaccuracies due to mechanical play. The repeatability of these linear axes typically lies between 1 and 0.1 lm. Some manufacturers and researchers claim that robots built with these axes reach an overall repeatability of 1 lm. A typical exponent of this class of robots is the Sysmelec Autoplace 411 (Fig. 17-2). Another solution for microassembly robots is conventional scara robots in combination with redundant high-precision axes in order to reach high resolution. This

17.3 Equipment Technology for Microassembly

Fig. 17-2 (a) Sysmelec Autoplace 411 serial robot and (b) RP5-AH parallel robot

approach is always combined with additional sensors in order to achieve good repeatability. The working platform of parallel robots is guided by several closed kinematic chains attached to the base frame, providing high structural stiffness. Owing to their light but rigid design and the ability to install all drives in a fixed frame, they are well suited for highly precise handling operations. A variety of parallel robots based on planar and spatial parallel structures with from three to six degrees of freedom (DOF) has been developed at the IWF so far. Three of these will be described exemplarily in the following.

17.3.2.1 MICABOE Planar Parallel Robot [10] This example considers a planar parallel structure (Fig. 17-3). Three linear drives move the platform with the gripper by three guiding chains equipped with CRP rods. The movement in the direction of the z-axis is performed by an additional elevating platform. The robot is driven by three piezoelectric stick– slip drives with smallest step sizes of 5 nm and is equipped with linear encoders with a resolution of 0.155 lm. First, the passive rotary joints were built up with conventional ball bearings (see also Section 17.5). The characteristics of the robot are summarized in Table 17-1.

17.3.2.2 TRIGLIDE Spatial Parallel Robot [11] Within the frame of a project sponsored by the German Ministry of Research and Education called ACCOMAT, the IWF and the company Robert Bosch developed an assembly cell for microassembly purposes. The fundamental element is a parallel robot which represents a spatial structure with three degrees of freedom in the x-, y- and z-directions (Fig. 17-4). The

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Fig. 17-3 MICABOE planar parallel robot Table 17-1 Characteristics of the MICABOE planar parallel robot

Structure

Performance data

Parallel robot with 3 DOF x, y, C + 1 DOF serial in z-direction Piezo drives (smallest step size 5 nm) Incremental encoders with 1 lm resolution Rotary joints with 1 DOF (ball bearings)

Velocity = 0.15 m/s Acceleration = 0.5 g Payload = 0.1 kg Repeatability = 2 lm Workspace = 30 ´ 30 ´ 30 mm/rotary angle 408 Footprint = 130 ´ 400 ´ 400 mm

three linear drives of the structure (TRIGLIDE) are arranged star-shaped in the base plane at intervals of 1208. Thus the structure has a workspace which is nearly triangle-shaped. The working platform is connected with each drive by two links forming a parallelogram, allowing only translational movements of the platform and keeping the platform parallel to the base plane. An additional rotary axis is serially integrated into the working platform. The orientation of the working platform is limited only by the gripper size and supply wires. This structure is very rigid and drive errors are reduced because the ratio of ‘platform movement’ to ‘drive movement’ is always < 1. In this configuration, the resolution of the electrical linear motors with linear encoders is 0.125 lm. The characteristics of the robot are summarized in Table 17-2.

17.3 Equipment Technology for Microassembly Fig. 17-4 Assembly cell with TRIGLIDE robot

Table 17-2 Characteristics of the TRIGLIDE spatial parallel robot

Structure

Performance data

Symmetric design (y-star) with 3 DOF x, y, z Hybrid structure with serial rotation axis C Parallelogram guidance of working platform Preloaded universal joints (2 DOF) Visual sensor guidance

Velocity = 0.2 m/s Acceleration = 5 g Payload = 1 kg Repeatability = 1 lm Workspace = 200 ´ 200 ´ 50 mm Footprint = 1280 ´ 980 mm

17.3.2.3 MICABOF Hybrid Parallel Robot [12] The MICABOF microassembly robot was developed especially within the context of the SFB 516 for the assembly of active microsystems. It has four degrees of freedom arranged in a hybrid robot structure (Fig. 17-5). Two linear axes with 0.1 lm resolution are used to drive the parallel structure in the x- and y-direction (x–y-plane). Inside the robot head, two serial mounted drives are located for motion in the z-direction and rotation around the z-axis. Furthermore, the robot head is designed hollow for the integration of a camera and a lighting unit. The characteristics of the robot are summarized in Table 17-3. 17.3.3 Pseudo-elastic Flexure Hinges in Parallel Robots

Other approaches to increase the accuracy of robots were made by using flexure hinges instead of conventional rotational joints, because backlash, friction and stick–slip effects in conventional joints decrease the overall precision of the robot. Particularly in parallel structures the high numbers of joints and joints with more than one degree of freedom are typical problems. The integration of flexure hinges in parallel structures is relatively simple, because except for the

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Fig. 17-5 MICABOF hybrid parallel robot

Table 17-3 Characteristics of the MICABOF hybrid parallel robot

Structure

Performance data

Symmetric design with 2 DOF x, y Hybrid structure with serial C-z-axis Air bearings Direct linear drives Visual sensor guidance

Velocity = 0.1 m/s Acceleration = 2 g Payload = 0.2 kg Repeatability = 2.6 lm Workspace = 120 ´ 100 ´ 15 mm Footprint = 480 ´ 600 ´ 280 mm

drives all joints are passive. Since flexure hinges gain their mobility exclusively by the elastic deformation of matter, they do not possess the above-named disadvantages of conventional joints. However, for the same reason their attainable angle of rotation is strongly limited when using conventional metallic materials with approximately 0.5% maximum elastic strains or plastics with *1% elastic strains. As a result, flexure hinges are mainly used in small devices such as micropositioning devices, microgrippers with small angular deflections and limited workspaces. Generally, a wide variety of different designs of flexure hinges are known. In most cases attempts are made to achieve high angular deflections with small oc-

17.3 Equipment Technology for Microassembly

curring elastic strains. Plastic deformations should be avoided owing to decreasing numbers of possible load cycles caused by fatigue of the material. In research projects at the IWF, different parallel robots were investigated in which the conventional bearings were replaced by pseudo-elastic flexure hinges. By using a shape memory alloy (SMA), which offers pseudo-elastic strains up to 17%, to build up flexure hinges, maximum deflection angles of ± 308 can be achieved. This offers the possibility of designing a robot with a sufficiently large workspace for typical microassembly tasks.

17.3.3.1 Pseudo-elastic Flexure Hinges Pseudo-elastic SMAs gain their reversible strains not only by distortion of the atomic lattice. Fig. 17-6 shows exemplarily a uniaxial stress–strain diagram of such a material. An SMA exists in at least two different phases (austenite and martensite) depending on the temperature or the applied stress. Initially a pseudo-elastic material is in its austenitic phase. At the beginning when loaded its deformations are elastic, with the conventional meaning. When further loaded, a transformation from the austenitic phase to the martensitic phase is induced. This transformation is accompanied by large reversible strains at nearly constant stresses. When unloaded, a retransformation from martensite to austenite occurs at lower stresses, resulting in a stress hysteresis. At the IWF, a single-crystal CuAlNi alloy with two stress plateaus and maximum reversible strains up to 17% is used. Since the material is now only available in rods, the design of the flexure hinges is restricted. Fig. 17-7 shows the optimized geometric dimensions of the flexure hinge used with one degree of freedom. In order to achieve high numbers of cycles, the maximum occurring strains were restricted to 4.2% at deflections of 308. Investigations of the kinematic behavior of flexure hinges showed that there are deviations between flexure hinges and ideal rotational joints. These deviations are caused by the movements of the instantaneous center of rotation during deflection. The deviations are also reflected in the kinematic behavior of the robot with integrated flexure hinges

Fig. 17-6 Stress–strain diagram of pseudo-elastic SMA

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17 Microassembly – Approaches to Meet the Requirements of Accuracy Fig. 17-7 Flexure hinge with one degree of freedom

(compliant mechanism) [12]. The error that is induced when describing the compliant mechanism by means of a rigid body model influences the absolute accuracy but not the resolution and the repeatability. In order to show the performance of flexure hinges, planar and spatial parallel robots were investigated analytically and experimentally.

17.3.3.2 Planar Compliant Robot (MICABOES) Fig. 17-8 shows the MICABOES planar robot with three degrees of freedom, generally described in Section 17.4, in which six rotational joints are replaced by six flexure hinges. With an assumed maximum angular deflection of ±308 at the hinges and maximum motion ranges of the drives of 70 mm, the workspace of the robot is 40 ´ 40 mm with no additional rotation of the platform. When using flexure hinges with conventional spring steel, the angular deflection is restricted to ±58. In this case the resulting workspace of the structure is about 20 times smaller than that using pseudo-elastic flexure hinges.

Fig. 17-8 MICABOES planar compliant parallel robot with three degrees of freedom

17.3 Equipment Technology for Microassembly

The mechanism was simulated by means of FEM. The simulation shows that the movement of the instantaneous center of rotation is also reflected in the behavior of the compliant mechanism which shows deviations from its rigid body model. The difference between the analytically computed tool center point (TCP) and the TCP of the compliant mechanism model exceeds 90 lm at certain points in the workspace. Starting from the undeformed initial position of the compliant mechanism, the deviations increase with the displacement of the TCP. Reasons are the increasing deflections of the hinges and the accompanying increasing deviations compared with ideal joints. The simulation indicates that flexure hinges principally reduce the absolute positioning accuracy but that the resulting deviations are comparably small. Additionally, the deviations of the single hinges are not accumulated but may compensate themselves.

17.3.3.3 Spatial Compliant Robot (TRIGLIDES) The spatial mechanism requires flexure hinges with more than one degree of freedom. Fig. 17-9a shows two possibilities for realizing flexure hinges with two degrees of freedom. In order to minimize the strains that occur, the smallest height h of the hinges (Fig. 17-7) has to be kept small. This will lead to a very small cross-sectional area in the first case and the hinge will be very sensitive to buckling and unintentional rotations. In the second case, there is an offset d between the two axes, leading to a different kinematic behavior. Therefore, flexure hinges with two degrees of freedom are realized by a spatial combination of flexure hinges with one degree of freedom. Fig. 17-9 b shows a first design of a combined flexure hinge with two degrees of freedom and intersecting axes replacing conventional universal joints. A problem with compliant mechanisms, especially spatial mechanisms, is their tendency to vibrate. In fact, the flexure hinges act as springs without any damping component, except for the inner damping of the deformed material. Fig. 17-10 shows an example of increasing stiffness and optimizing the distribution of occurring forces by a suitable design of a combined flexure hinge. Owing to the parallel and angular arrangement of the hinges, torsional moments can be absorbed and transformed into tension and compression forces better.

Axis 2

Axis 1

Fig. 17-9 Flexure hinges with two degrees of freedom

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Fig. 17-10 Optimized combined flexure hinge

Fig. 17-11 TRIGLIDES compliant spatial robot with three degrees of freedom

Fig. 17-11 shows a compliant spatial robot with three degrees of freedom and six integrated combined flexure hinges. The structure is generally described in Section 17.4. Restricting the deflection angle of the hinges to ± 308, the workspace is only minimally reduced compared with that with no angular restriction. With the actual configuration, a cube with dimensions 112 ´ 112 ´ 112 mm fits into the workspace. Table 17-4 shows the dependence between the achievable workspace and different restrictions of the angular deflection of the hinges which depend on the material used. Using flexure hinges made of spring steel, the resulting workspace is 100 times smaller than that when using pseudo-elastic flexure hinges. To verify the kinematic behavior of the compliant mechanism, the same FEM investigations as for the planar mechanism were performed. A planar movement in an area of 60 ´ 60 mm leads to maximum deviations of *1.5 mm and for movement of 60 mm in the z-direction the deviation is *0.05 mm. Consid-

17.3 Equipment Technology for Microassembly Table 17-4 Workspace with different angle restrictions

Angular deflection (8)

x (mm), min., max.

y (mm), min., max.

z (mm), min., max.

Workspace (normalized)

± 1808 ± 308 ± 58

–218, 176 –202, 176 –34, 34

–206, 206 –174, 174 –30, 30

–324, 0 –314, –26 –220,–170

1 0.88 0.008

ering the high number of flexure hinges, this deviation is relatively small but not categorically negligible. In the literature, approaches for increasing the absolute accuracy of a compliant robot exist. This is done, for example, by adding additional actuators or by varying the link lengths depending on the deflection of the hinges. In order to use these approaches, a calibration of the mechanism is required. For example, a difference of 0.1 mm of one link length will lead to deviations of *65 lm at the planar robot. This error is of the same range as the deviation caused by flexure hinges. Depending on the application, it has to be decided whether the effort to compensate for the deviation is necessary. For many assembly tasks, a high resolution and repeatability should be sufficient [13], which are not influenced by these kinematic deviations.

17.3.3.4 Experimental Results The performance of robots was investigated by measuring the repeatability (RP) and additionally the variance of multiple direction position accuracy (vAP) for the spatial robot. Capacitive sensors with a resolution of 8 nm were used and the measures and calculations were carried out according to EN ISO 9283. The results (average values) are given in Table 17-5. Both compliant robots, especially the spatial robot, achieve very good repeatability and resolution owing to the lack of backlash and friction. Compared with the values for robots with conventional joints, it was possible to increase the accuracy more than twofold. Approaching the same desired pose from three directions orthogonal to each other (measuring the vAP), backlash in joints plays a major role and for conventional robots vAP is often much worse than RP. With vAP = 0.15 lm at the TRIGLIDES an outstanding performance of the robot was achieved. Table 17-5 Measured repeatability of the compliant robots

Robot

Repeatability at 3r

Workspace used (mm)

MICABOES TRIGLIDES

1.0 lm (x–y-plane) 0.3 lm spatial 0.2 lm (x–y-plane) 0.15 lm (vAP)

22 ´ 22 30 ´ 30 ´ 30

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Owing to the compliance, the settling time to reach a stable position is higher than with conventional joints, e.g. 0.7 s for the planar compliant robot compared with 0.15 s for the conventional robot. Vibrations are particularly excited by stick–slip effects of the piezo drives. For this reason, drives with no inherent stick–slip effects should be used for compliant mechanisms. For the spatial robot, vibrations could be minimized because of the optimized design of the structure and the combined flexure hinges. The settling time at this robot is determined only by the control parameters. There is no measurable difference between the conventional structure and the structure with flexure hinges. 17.3.4 Grippers for Microassembly

During the handling of smallest parts with microgrippers, problems may arise when the surface forces, such as van der Waals, electrostatic or capillary forces become larger than the weight. In this case, the components cling to the gripper and can only be separated from it with difficulty. Several investigations and theoretical approaches referring to this problem can be found in the literature, because the problems arising are very application specific. In general, grippers can be arranged into three groups depending on the gripping principle: · by use of the adhesive principle; · by use of the positive gripping principle; · by use of the non-positive gripping principle. Since grippers using the adhesive principle may contaminate the gripped parts and cannot be miniaturized owing to their complex construction, this type of gripper is not commonly used in microassembly. The positive and non-positive gripping principles are often found in microassembly. Once again they can be sub-classified into mechanical, vacuum and electrostatic grippers. A wide range of these types of grippers for the assembly of miniature parts have been developed and are already commercially available in research and industry. As the gripper represents the direct interface between the part and the assembly robot, a careful choice is essential for a successful microassembly process. In the case of an optical sensor guidance of the assembly process, the basic requirement is the possibility of observing the grasped part with an optical sensor, mostly fixed to the tool head of the robot. This can be achieved by using either partly hidden or transparent grippers or jaws. In the scope of the SFB 516, two different gripping principles have been closely investigated and some prototypes have been manufactured (Fig. 17-12).

17.3.4.1 Mechanical Gripper A modular miniature gripper made of polyoxymethylene (POM), featuring flexure hinges, a shape memory actuator and exchangeable jaws, has been devel-

17.4 Sensor Guidance in Microassembly

Fig. 17-12 (a) Mechanical microgripper and (b) electrostatic gripper

oped at the IWF [14]. In a joint effort together with the Institute for Microtechnology (IMT), Braunschweig, it was possible to miniaturize this gripper principle even further by means of microtechnology. It is driven either by shape memory actuators [15] or by a newly developed pneumatic actuator for use in conjunction with microgrippers [16].

17.3.4.2 Electrostatic Gripper In contrast to the electrostatically supported manipulation of particles in the vacuum chamber of a scanning electron microscope (SEM), the centering electrostatic microgrippers operate under atmospheric conditions. The gripper grasps and centers parts by means of a non-uniform electrostatic field [17]. The field is generated through voltage supply to a concentric set-up of an inner and an outer electrode. The gold electrodes are structured on a transparent Pyrex substrate. This gripper is suitable for the handling of very small parts such as spheres of a microball bearing in particular.

17.4 Sensor Guidance in Microassembly

Since in assembly processes the positioning failures of the handling systems are superimposed by manufacturing and positioning tolerances (e.g. in the magazine) of the single components and by external effects such as vibrations or thermal effects, sensors are more and more used for sensor-controlled assembly systems in addition to high-precision positioning equipment. For this purpose, optical sensors with a commercial or self-developed image processing system are used. Additionally, force measuring can be used to sense the contact between assembly partners in the z-direction or to limit the insertion forces. For

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this reason, conventional assembly robots are equipped with unidirectional force sensors. Common performance data are a resolution of 1.5 mN with a minimum load of 10 mN (e.g. Sysmelec Autoplace 411). To sense hidden edges in the x–y-plane, a three-dimensional force sensor is needed, but is still not stateof-the-art. Sensor guidance for automated assembly of microsystems is mostly based on the visual sensing of tags or objects. The objective is to calculate a relative position vector between the parts that have to be mated. Within the work of the SFB 516, a 3D vision sensor using only one camera was developed at the Institute of Production Measurement Engineering (IPROM), Braunschweig. It uses the principle of stereo photogrammetry. With a special arrangement of a reflecting prism and two flat mirrors, two different perspective views of the object are obtained. A CCD camera with 768 ´ 568 pixels (384 ´ 568 pixels for a half frame) is used. The image area of 7.5 ´ 11 mm is defined by the focal length of the lenses (12 mm). For this field of view a resolution of *19 lm per pixel is obtained. It varies with the distance from the object to the 3D sensor. The measurement repeatability of the 3D vision sensor was measured against a laser interferometer with 0.3 lm in the x–y-plane and 0.9 lm in the z-direction. This 3D vision sensor was integrated into the robot head of the MICABOF. An interface between the developed vision system and the robot control system permits a 3D vision guided microassembly. 17.4.1 Microassembly Scenarios

In the following, the assembly of different microsystems with special challenges during the assembly process is considered. As different assembly systems are used for the scenarios, the assembly systems are difficult to compare and only the advantages and disadvantages in the assembly scenarios are described. The main difference between the assembly systems used is the concept of sensor guidance. For position correction in the x–y-plane and rotation the Sysmelec Autoplace 411 uses a 2D vision sensor. Placing of the part (in the z-direction) is realized with a force sensor. In contrast, the MICABOF uses the 3D vision sensor for position correction in the x–y–z-direction and rotation.

17.4.1.1 Assembly Scenario I (Sysmelec Autoplace 411) In the first assembly scenario, the Sysmelec Autoplace 411 is used to assemble a horizontal microelectric linear motor (Fig. 17-13) developed within the SFB 516 [18]. The linear motor works according to the reluctance principle. It consists of a ceramic stator with six coil systems meandering horizontally around soft magnetic poles, a glass traveler with soft magnetic poles corresponding to those of the stator and a ball bearing in between. The stator is fixed in a silicon die which has two V-grooves used as ball tracks. Ruby balls are placed inside

17.4 Sensor Guidance in Microassembly

Fig. 17-13 (a) Assembly scenario I and (b) horizontal microelectric linear motor

the V-grooves and form the ball bearing by carrying the weight of the traveler. Fig. 17-15 shows the assembly order. For these assembly tasks, a vacuum gripper is used. The first challenge in the assembly is placing the balls (d = 0.6 mm) in the V-groove. If only the vacuum is switched off, the ball remains on the nozzle of the gripper because of remaining negative pressure and adhesive forces. The difficulty is to pulse an air stream with an adequate pressure through the nozzle which is just sufficient to disengage the ball. If the pulse is too strong, the ball will be blown away. This process has to be conducted six times, because there are three balls in each Vgroove. By placing the second or third ball in the V-groove, the already placed ball(s) can be blown away by the pulse of the nozzle. Therefore, placing of the balls starts on the side of the V-grooves that point away from the inclined nozzle, because the pulse will not directly affect the placed balls. Another critical assembly task is the positioning of the traveler above the stator and the insertion of the passive pole rows of the traveler between the active pole rows of the stator. The horizontal gap between the pole teeth amounts 3 lm over a length of *6 mm. It is obvious that a high assembly precision is necessary to fulfil these requirements. Therefore, positioning of the traveler is done with 2D visual sensor guidance. A special challenge is to identify properly a pattern on the object. In addition to stable lighting, an exact camera calibration of the scene is important. Unfortunately, the calibration accuracy of the vision system is not better than 1 lm. Therefore, 1 lm is the best reachable accuracy with this assembly system. First, the position of the stator is measured and saved in the control system. Subsequently, the position of the traveler is measured. After position correction in the x–y-direction and rotation around the z-axis, the traveler is moved to a

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calculated position over the stator. This position is calculated, resulting in the measured position of the stator and the corrected position of the traveler. After this movement, no further measurement is made, because the stator is hidden behind the traveler and no pattern can be found. As a result of the last movement without any further optical measurements the positioning accuracy is in the range of the robot’s repeatability. When the traveler is positioned above the stator, it will be inserted between the active stator poles by a movement in the z-direction. To prevent destruction of the system, a force sensor is used, which stops the movement when a limit value is reached. In this position, the vacuum gripper releases the traveler, the robot moves upwards and the assembly is finished.

17.4.1.2 Assembly Scenario II (MICABOF) The second assembly scenario shows the positioning of guides for a vertical microelectric linear motor developed within the SFB 516 [19]. This motor consists of a stator with active poles, a traveler with active poles and two linear guides (Fig. 17-14 a). The main difference between this motor and the horizontal microelectric linear motor is the effective direction of the coil. Fig. 17-16 shows the assembly order of this scenario. Although the assembly steps are equal for the right and left guides, it is only pictured up to the point where the assembly of the left guide starts. For the positioning of the guides, the self-developed assembly system with 3D sensor guidance is used (Fig. 17-14 b). Each element of the vertical microelectric linear motor is provided with specific marks which can be recognized by the sensor. In a first step, the marks on the stator are checked. If the part can be recognized, the positioning of guides can be carried out. Otherwise, the checked stator will be omitted. The guides are also checked, and are picked up if they are recognized. Picking up the guide and moving it into a predefined position over the stator is executed without sensor guidance. Subsequently the sensor guidance is started automatically. It oper-

Fig. 17-14 (a) Vertical microelectric linear motor and (b) assembly scenario II

17.4 Sensor Guidance in Microassembly Fig. 17-15 Assembly order of horizontal microelectric linear motor

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ates in a look-and-move procedure. The vision system measures the marks of stator and guide, calculates their distance and gives a relative positioning vector to the robot control. Inside the robot control the relative position vector is separated into a rotation correction, a correction in the x–y-direction and, finally, the correction in the z-direction which places the guide. Each correction step has a break criterion in the form of a limit value. If the rotation correction is below its limit value, the correction in the x–y-direction will start and if this is below its limit values the guide will be placed. Owing to the limit value for a x–y-correction of 2 lm (radius around the target position), the reachable positioning accuracy is below this value. Fixing the guides is still an unsolved problem in this assembly task. Therefore, research in the field of microgluing will take place in the SFB 516. Currently, the positioning accuracy achieved can only be measured by the 3D vision sensor before opening the gripper. When the gripper opens, the part is not re-

17.5 Conclusion

producibly shifted out of the placing position because of the adhesive effect and the missing fixture.

17.4.1.3 Comparison of the Microassembly Scenarios For both assembly scenarios, sensor guidance is used, but with different concepts. In the first scenario, sensor guidance is implemented through a combination of a 2D vision sensor and a force sensor. After a precise positioning (in the x–y -direction and rotation around the z-axis) of one part relative to another with the 2D vision sensor it is placed by using a force sensor while driving in the zdirection. If a limit value is reached, the movement will stop and the vacuum gripper will release the part. With this concept, a defined joining force can be guaranteed. For the second scenario, a 3D vision sensor is used. With this concept, distances between marks on the parts can be measured in the x-, y- and z-direction. As each part is equipped with at least two marks, rotation of the parts is calculated by their measured coordinates. The accuracy of placement in the z-direction depends on the measurement repeatability of the 3D vision sensor in the z-direction and the compliance of the real part height from the desired part height. Therefore, the joining force cannot be guaranteed with this concept. Placing of parts is done in one robot movement and is faster than with a force sensor. This method is only feasible if the joined parts are not too fragile. As a future development, the integration of a force sensor in the self-developed assembly system is planned. It should be used for assembly scenarios of parts with high sensitivity whereas in other cases the presented concept with the 3D vision sensor should be used. In due course, the sensor guidance concept will be changed towards a continuous measurement and position correction during robot movement to decrease the assembly time.

17.5 Conclusion

Depending on the application areas, the variety of aspects related to microassembly is wide. Many authors mention bare die handling in electronic production or the assembly of optical components by talking about microassembly. This chapter gives an insight into aspects of microassembly regarding the concrete example of active microsystems. The work and the developments were mainly motivated by the assembly tasks required for the manufacture of active microsystems in the SFB 516. The chapter has focused on equipment technology and cannot cover the joining technology, which is a challenging research area itself. In summary, the involved and important accuracies, robots available for microassembly tasks and their respective accuracies are presented here. Additionally, grippers and sensors available for a sensor-guided microassembly system are discussed.

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17.6 References 1 F. Götz (ed.), Nexus MST Market Analysis

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2002; Grenoble: Nexus Association, 2002. 2. Öffentliches Statusseminar zum Verbundprojekt, Microfeed2: Produktstandards und Fertigungsmittel für das Vereinzeln, Ordnen und Magazinieren von Mikrokörpern und Mikrosystemen, gefördert durch das BMBF, 1998–2001. S. Koelemeijer Chollet, J. Jacot, Cost efficient assembly of microsystems. mstNews 1999, January, 30–33. Fertigungsmittel für Mikrosysteme – Werkstückträger – Anschlussmaße und Toleranzen; DIN 32561, April 2000. Arbeitsgemeinschaft Match-X – Modulare Mikrosysteme; http:// www.pb.izm.fhg.de/match-x/index.html. A. H. Slocum, Precision Engineering Research Group; http://pergatory.mit.edu, 2004. Industrieroboter: Leistungskenngrößen und zugehörige Prüfmethoden; EN ISO 9283, Berlin, 1999. Montage- und Handhabungstechnik, Kenngrößen für Industrieroboter; VDI 2861, Blatt 1–3, Düsseldorf: VDI, 1988. J. Hesselbach, A. Raatz (eds.), MikroPRO – Untersuchung zum internationalen Stand der Mikroproduktionstechnik; Essen: Vulkan, 2002. J. Hesselbach, G. Pokar, A class of new robots for microassembly. Prod. Eng. 2000, 7, 113–116. J. Hesselbach, J. Wrege, A. Raatz, O. Becker, Aspects on design of high precision parallel robots. J. Assembly Autom. 2004, 24, 49–57.

12 J. Hesselbach, A. Raatz, Pseudo-elastic

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flexure-hinges in robots for microassembly. Proc. SPIE 2000, 4194, 157–167. P. Schellekens, N. Rosielle, H. Vermeulen, M. Vermeulen, S. Wetzels, W. Pril, Design for precision: current status and trends. Ann. CIRP 1998, 47, 557–586. R. Pittschellis, Mechanische Miniaturgreifer mit Formgedächtnisantrieb; Fortschritts-Berichte No. 714, Reihe 8, Düsseldorf: VDI, 1998. S. Bütefisch, S. Büttgenbach, G. Pokar, J. Hesselbach, A new SMA actuated miniature silicon gripper for micro assembly. In: Proc. 7th Int. Conf. on New Actuators, Bremen; 2000, pp. 334–337. S. Bütefisch, S. Büttgenbach, A new pneumatically actuated miniature gripper for micro assembly. Proc. SPIE, 2001, 4568, 32–39. J. Hesselbach, J. Wrege, C. Graf, S. Büttgenbach, S. Bütefisch, Centering electrostatic microgripper and magazines for microassembly tasks. Proc. SPIE 2001, 4568, 270–277. V. Seidemann, J. Edler, S. Büttgenbach, H.-D. Stölting, Linear variable reluctance (VR) micro motor with horizontal flux guidance: concept, simulation, fabrication and test. In: 12th International Conference on Solid-state Sensors, Actuators and Microsystems, Transducers 2003, Boston; 2003, pp. 1415–1418. M. Föhse, J. Edler, H.-D. Stölting, H.-H. Gatzen, Design, fabrication and characterization of a miniature linear asynchronous motor. In: Proc. Actuator 2002, Bremen; 2002, pp. 176–179.

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18 Quality Assurance and Dimensional Measurement Technology J. Fleischer, I. Behrens, Institute of Production Science (wbk), University of Karlsruhe (TH), Germany Abstract The need to develop new concepts and methods in microproduction is in particular valid for the development- and production-accompanying process of quality assurance owing to the challenges that specifically arise regarding the application of quality management methods and dimensional micrometrology. In this context, the specific approach of the subproject C4 related to quality assurance within the Collaborative Research Centre (SFB) 499 ‘Design, Production and Quality Assurance of Primary Shaped Micro Components Manufactured in Ceramic and Metallic Materials’ is discussed. One core aspect is the development of the microfeature concept in order to collect, store and provide information throughout the whole microproduction process chain. Another important issue is the analysis of quality in dimensional micrometrology.

Keywords quality assurance; microfeature; dimensional micrometrology

18.1 18.2 18.2.1 18.2.2 18.3 18.3.1 18.3.2

Introduction 500 Quality Assurance for the Production of Micromechanical Parts 500 Challenges of Applying Methods of Quality Assurance 501

Challenges of Measuring Micromechanical Parts Dimensionally 501 State-of-the-Art 501

State-of-the-Art of Methods of Quality Assurance 502 State-of-the-Art of Measuring Micromechanical Parts Dimensionally 503

18.4

Approaches to Quality Assurance and Dimensional Metrology in the SFB 499 507

18.4.1 18.4.1.1

Methods and Quality-related IT Support in the SFB 499 507 The Microfeature 508

Advanced Micro and Nanosystems Vol. 4. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31493-8

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18.4.1.2 18.4.2 18.4.2.1 18.4.2.2 18.5 18.6 18.7

The Microfeature Catalog 509 Dimensional Metrology in the SFB 499 512 Measurement Results 513 Measurement Uncertainty 514 Conclusions and Outlook Acknowledgment 517 References 518

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18.1 Introduction

The design and production of wear-resistant micromechanical components requires the development and optimization of a wide range of both manufacturing and peripheral processes, as can be seen in Sections 2 and 3 and Chapter 16. New concepts and approaches have to be developed, analyzed, optimized and validated in order to fulfil the prerequisites for a successful transfer of these new technologies into industrial application. The need to develop new concepts and methods is in particular valid for the development- and production-accompanying process of quality assurance. Quality assurance is essential for the efficient support of the manufacturing and peripheral processes from design through production to automation with the major objective of achieving robust processes with a high throughput of microparts meeting their functional requirements. The following two sections outline the challenges and the state of the art of quality assurance and in particular dimensional measurement technology for micromechanical parts. The specific approach of the research project C4 related to quality assurance within the SFB 499 is then discussed.

18.2 Quality Assurance for the Production of Micromechanical Parts

Quality assurance as a whole comprises methods and approaches which are either preventive or production-accompanying. The preventive methods include FMEA (Failure Modes and Effects Analysis), FTA (Fault Tree Analysis) and QFD (Quality Function Deployment). Preventive methods are primarily employed during the conception, development and process planning phases within a product life cycle. The objective is to locate risks and thus to assure the quality of the final product at an early stage. Complementarily, the methods for production-accompanying quality assurance such as SPC (Statistical Process Control) are applied mainly during the production phase [1]. For all these methods and approaches, the crucial precondition for a successful application is a comprehensive set of data and information on the product and the processes, the functional interrelations of the microsystem, the influ-

18.2 Quality Assurance for the Production of Micromechanical Parts

ence of process parameters on the quality of the product, etc. Apart from the analysis of material characteristics and the experience gained during production, information on the compliance with dimensional, shape and position tolerances of microparts is essential not only in order to assess the quality of microparts but also in order to monitor and optimize production processes. According to Pfeifer et al. [2], among all the tasks of measurement material testing, integrity inspection and tests on functionality account for only 10% of all metrological requirements, whereas 90% are constituted by dimensional measurements. Even more, the dimensional measurement is anything but trivial in the micrometer and sub-micrometer range. Therefore, the dimensional measurement of microparts is treated in detail in this chapter. 18.2.1 Challenges of Applying Methods of Quality Assurance

To achieve the economically efficient production of micromechanical parts through high-quality and robust production processes, the application of suitable quality-assuring methods accompanying both the development and the production is essential. However, it is not possible to directly apply those methods which are commonly used for products in macrodimensions. One reason is that a micromechanical part is almost completely composed of functional surfaces. Moreover, the complex interdependences between different manufacturing steps due to the dimension of the structures have to be taken into account. In addition, no experiences with respect to microdimensions initially existed for the production and the corresponding accompanying processes. Consequently, the existing methods of quality assurance have to be adapted or even combined to be successfully implemented in microproduction processes. This especially includes the fact that concepts and structures have to be defined and implemented, which guarantee the necessary exchange of information between all participating subprocesses in order to close the information gap as quickly and in as user-friendly a manner as possible. Accordingly, quality assurance has to focus first on the collection, structuring and supply of processand product-related data. 18.2.2 Challenges of Measuring Micromechanical Parts Dimensionally

There is a wide range of different micromechanical parts with regard to their geometric, surface and material characteristics. Consequently, the number of different measuring tasks is manifold. The structural details of micromechanical parts are smaller than 1 mm, they are typically two and a half- or three-dimensionally structured and they have high aspect ratios (10–100). Moreover, edges are often rounded or have a chamfer. The dimensional, shape and position tolerances are within the micrometer or even sub-micrometer range. The reflection of surfaces of micromechanical parts, which is in particular relevant

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for optical measurement techniques, differs significantly. Roughness and waviness of the surfaces are also diverse. Concerning the material characteristics, micromechanical parts are either made of plastics, metal or ceramic. Based on these characteristics of micromechanical parts, the following challenges for dimensional measurement arise. First, requirements regarding the precision of a measurement are very high owing to tolerances down to 1 lm and below. Second, the measurement device has to be capable of detecting and measuring three-dimensional structural details in the micrometer range. Third, measurements have to be carried out in such a way that the micromechanical component is not destroyed, damaged or influenced. Fourth, dimensional metrology faces the challenge to accurately measure objects with various possible surface characteristics without adversely influencing the precision of measurement. These challenges resulting from the characteristics of micromechanical parts are complemented by those resulting from the production of micromechanical parts. Thus, in order to economically assure the product quality and to achieve and control robust production processes, the measurement not only has to be very accurate but also fast and close to the production line. Only this permits a quick feedback loop to control the production efficiently.

18.3 State-of-the-Art 18.3.1 State-of-the-Art of Methods of Quality Assurance

There are several internationally valid documents and guidelines describing methods and standards of quality assurance for macroscopic parts, e.g. ISO 9000 ff. Methods such as QFD and FMEA have been successfully introduced and implemented in order to improve the quality of products and processes. In contrast, only a few approaches regarding the methodical quality-related support of microproduction processes are known. For micropowder injection molding (microPIM) processes, two relevant holistic methodical approaches for the quality management of this production process covering the whole process chain are known [3, 4]. In both cases the research was subject to the introduction of a standardized quality management system according to DIN/ISO. In the course of the FEPAKOM (production technology for passive optical components) research project [3], a methodical approach to certify a company in the field of serial microproduction was proposed. In another project, a concept for the ISO certification of a research institute was developed [4]. In particular, the implementation of preventive quality management methods is recommended. However, the explicit adaptation of methods for quality assurance in the production of micromechanical parts has not yet been carried out.

18.3 State-of-the-Art

In addition, research work in the field of process simulation and modeling, which indirectly contribute to quality assurance, is known. These research activities focus on the optimization of production processes and the corresponding process parameters [5, 6]. Both in industrial application and in research, the predominant issues related to quality in microproduction are process surveillance and process visualization [7, 8] and the development of measurement technology. The reason is that adequate devices for the measurement of quality assurance data for both process and part are not yet readily available in microdimensions. 18.3.2 State-of-the-Art of Measuring Micromechanical Parts Dimensionally

In order to collect information on the geometry of microparts, there basically exist the following possibilities: scanning electron microscopy, scanning probe microscopy, optical and tactile methods (see Fig. 18-1) Fig. 18-2 gives an overview of resolutions and measurement ranges. In the following, the potential of these methods is discussed with regard to their efficient applicability for quality assurance in microproduction. Scanning electron microscopy (SEM) provides both a high magnification (up to 5 ´ 105 ´) and a high depth of focus with resolutions down to 0.001 lm [10]. Further, the images are rich in contrast. Hence SEM technology allows one to generate detailed images of three-dimensional microstructures. Dimensional information can only be withdrawn by manually placing a scale within the picture. Even though the edge sharpness is high, the manual process makes measurement results uncertain. Even more, SEM measurements take place in an evacuated environment, which makes the measurement process complex and time consuming. Therefore, SEM is not suitable for an application with quick feedback loops into the production line. As a result, SEM technology is mainly

Fig. 18-1 Overview of possible measuring methods

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Fig. 18-2 Resolution and measurement ranges [9]

used as a versatile tool for the characterization of materials, but it is not adequate for the dimensional measurement of micromechanical parts. Scanning probe microscopy (SPM) comprises several different methods (see Fig. 18-1), which all work according to the same principle. The surface of the object is scanned with atomic resolution in order to obtain information on the surface characteristics. The prevalent methods are atomic force microscopy (AFM), which is a synonym for scanning force microscopy (SFM), and scanning tunneling microscopy [11]. Apart from the sub-ångstrom resolution, a magnification up to 107 ´ is feasible, for example, with the AFM technology. In contrast to SEM measurements, a time-consuming preparation and the application of vacuum are not necessary. Still, SPM measurements are time intensive owing to the sequential scanning process. Moreover, it is impossible to probe 2.5D or even 3D microparts. For example the walls of a channel or the geometry of a hole cannot be measured by SPM. There is also a wide range of optical measurement methods from focus techniques through interferometric methods to triangulation. With regard to the dimensional measurement of micromechanical parts, the main advantage of optical methods is the contact-free measurement, which makes these methods particularly applicable for fragile objects. Nonetheless, the precision of an optical measurement depends heavily on edge sharpness and on the illumination and consequently on the optical characteristics of the micropart. Therefore, a highprecision optical measurement is impossible in many cases. Moreover, the geometric measurement of rounded edges and high aspect ratios poses problems for optical methods.

18.3 State-of-the-Art

In tactile methods, the geometry of objects is measured by touching the object with a specific sensor. Tactile spherical probes in classical coordinate measurement technology in an example. The advantage of tactile sensors is their 3D capability. Moreover, tactile probes with scanning capability allow quick measurements. On the negative side, the main disadvantage is the minimum diameter of a probing sphere of 200 lm, limiting the resolution. The active forces on the measured object are another drawback. Overall, none of these methods fully meets the requirements of the dimensional measurement of micromechanical parts either with respect to a fully developed three-dimensional capability or with respect to the required precision. Therefore, sensorial concepts for the dimensional measurement of microparts have been developed in the past [12]. Owing to the advantage of tactile probing, especially hybrid solutions combining tactile and other measuring methods have been designed. Of special relevance are the Werth fiber probe (WFP) of Werth Messtechnik (Fig. 18-3 a), which combines optical and tactile probing, and the UMAP-130 of Mitutoyo (Fig. 18-3 b), which is based on an oscillation principle. The WFP is a glass-fiber with a melted spherical tip as probing element. The tip is positioned at the focal point of the optical system. When probing the measured object, a CCD camera takes images of the tip and the position of the tip is analyzed with an image processor [13, 14]. Regarding the UMAP-130, the probing element oscillates at a specific frequency. As soon as the probing element approaches the object, the amplitude changes. The probing element is moved towards the object until a specific reduction of the amplitude corresponding to a relative distance of the probing element from the object is reached [15]. The dimensions of the probing element are similar for both commercially available sensors with a minimum diameter of 25 lm of the WFP [16] and a diameter of 30 lm of the UMAP-130 [17]. The WFP is a 2.5D sensor, whereas the UMAP-130 is a 3D sensor, but the precision is better for the WFP. For both sensors the probing forces are below 1 lN, which is sufficiently small [16, 17].

Fig. 18-3 (a) Werth fiber probe (Werth Messtechnik) [14] and (b) UMAP-130 (Mitutoyo) [13]

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Recently, other microsensors have also been developed. For example, the WFP sensor has been extended to the third dimension by a second camera which focuses an additional target mark on the shaft of the glass-fiber from the side (Fig. 18-4 a) [18]. Two additional 2D sensors have been developed and prototypes constructed at the University of Tokyo. One of them is an opto-tactile sensor, where the measured object is touched with a spherical metal probe, on which a laser beam is focused (Fig. 18-4 b). Based on the deflection of the reflected beam which is absorbed by a photodiode, the probed point is evaluated. The diameter of the prototype is 5 mm [19]. The other sensor is based on a pneumatic-tactile principle. A probing ball with a diameter of 100 lm is placed below an evacuated pipe. When touching an object, the ball is displaced, air floats into the pipe and the pressure difference is analyzed (Fig. 18-4 c). The probing forces are below 1 lN [20]. Finally, an electro-tactile 3D sensor has been developed based on silicon-membrane-spring technology. The displacement of

Fig. 18-4 (a) 3D fiber probe [22], (b) opto-tactile sensor [19], (c) pneumatic-tactile sensor [20] and (d) electro-tactile sensor [23]

18.4 Approaches to Quality Assurance and Dimensional Metrology in the SFB 499

the probe with a diameter of currently 110 lm is detected and evaluated via piezoresistors (Fig. 18-4 d). It is possible to detect displacements in all translatorial axes independently [21]. To sum up, the new sensorial concepts offer the highest potential concerning the high-precision measurement of three-dimensional micromechanical parts. However, no fully capable solution to meet all requirements of micrometrology has yet been developed.

18.4 Approaches to Quality Assurance and Dimensional Metrology in the SFB 499

The prevailing issue for subproject C4 during the research performed so far has been to create the boundary conditions for the successful adaptation and application of quality assurance methods. These essential boundary conditions are first the development of structures for information collection and transfer and for cooperation. Second, the establishment of an adequate micrometrology for micromechanical parts given the restrictions of the available measuring technology is crucial to successfully support the other subprocesses involved in the development of robust microproduction processes. 18.4.1 Methods and Quality-related IT Support in the SFB 499

Experience has shown that a manufacturing-oriented construction and a robust molding of high-strength, wear-resistant microparts is possible only if a constant exchange of information between the subprocesses of the microPIM or the microcasting process including quality assurance takes place. However, successful interdisciplinary cooperation is impeded by lack of experience and by a temporally and locally distributed generation of information. Consequently, structures for ‘documented’ experience and the know-how about the historical development of process parameters as a decisive factor for the fabrication of high-quality microparts have to be created. Results and analyses of semi-finished and finished microproducts have to be saved in an appropriate way. They must be provided together with coherent information throughout the entire process chain. As a means for structuring relevant information, the feature concept was chosen and adapted to microdimensions. This choice was made because the feature concept has already been successfully applied for the optimization of macroproduction processes by providing defined additional information regarding the process flow. Based on the specifically adapted feature concept, an Internetbased application, the microfeature catalog, was developed to guarantee a platform-independent distributed collection and provision of information.

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18.4.1.1 The Microfeature The first feature concepts date back to the 1970s. Features were first defined by Grayer [24] in 1976 as a geometric area that can be generated by a single machine operation. Since the middle of the 1980s, features have been increasingly applied not only in research but also in commercial computer-aided design (CAD) tools. Further developed features were mainly integrated in the design process. Today, some commercial CAD systems offer feature catalogs related to the know-how of macromilling and -drilling operations in order to accelerate the design process. Regarding the current state-of-the-art, features are known as fabrication, quality or design features [25]. Correlated to defined phases of the product life cycle, these features constitute a collection of specific information and information regarding a product (e.g. geometry, technology, fabrication data). In comparison with the mere geometric description, they represent elements on a semantically higher level, which can be applied by different departments of an enterprise [26]. The definition of microfeatures for micromechanical components developed at the Institute of Production Science overcomes the strict separation of the semantic aspects of a geometric component area into categories of characteristics and aspects that are specific for the product life cycle. Microfeatures present a holistic approach that corresponds to the specifics and development demands of micromechanical parts with a particular focus on molding processes [27] in order to efficiently leverage the advantages of the implementation of feature concepts. The microfeature is defined by means of a geometric functional element which is determined by functional requirements within the design process. The core part of the microfeature definition is constituted by the nominal values of geometric dimensions including dimensional, shape and position tolerances and the specified surface quality. Moreover, fabrication-specific process information is incorporated into a microfeature. This is realized by attaching all relevant process parameters to a microfeature. In the application at hand, these are the process parameters of all processes involved from the production of the molds through molding to debinding and sintering. Owing to the lack of process-related experience and know-how, initially all process parameters are relevant. Additionally, the production-dependent constraints of microproduction processes are incorporated into a microfeature in addition to information regarding the material (for example, feedstock data) and its properties. Furthermore, detailed metrological information is also integrated into a microfeature in order to assure the comparativeness in particular with regard to dimensional measurements. Fig. 18-5 summarizes these five components of this microfeature definition. Microfeatures are methodically deduced by applying axiomatic design, the decomposing structuring method originally invented by Suh [28]. To deduce microfeatures in general and then apply them to a specific set of micromechanical parts, this method had to be adapted to the micrometer regime. As a result, microfeatures as integrating process information objects constitute knowledge carriers along the whole process chain. The holistic approach of

18.4 Approaches to Quality Assurance and Dimensional Metrology in the SFB 499

Fig. 18-5 Definition of a microfeature [27]

the microfeatures is of decisive significance for their application. Only with the continuous and process-comprehensive application along the process chain of the product development cycle can the total potential of benefits evolve. In this context, microfeatures implement the recording and classification of acquired process knowledge and the process experience. Furthermore, they facilitate the well-directed support of the development and design of microcomponents by indicating critical microfeatures from the manufacturing point of view.

18.4.1.2 The Microfeature Catalog In order to store and provide the microfeature-structured information an information platform, the microfeature catalog, has been developed and established exemplarily for the molded demonstrator objects (microgear and microturbine) of the SFB 499 and the production processes at hand. The microfeature catalog is based on a Microsoft SQL 2000 database. The interface to the user is implemented as web pages so that there is no need for specified software apart from Internet Explorer. The underlying relational database was systematically developed by constructing an entity relationship model of the microfeature-structured data.

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Fig. 18-6 Realizations of the microfeature ‘single-stage hole’

To increase user-friendliness, the microfeatures themselves are classified into ‘families’ on the highest level of the catalog. The clustering is accomplished following macromilling and -drilling features, which include about 80% of the processed objects in the conventional macro range [29]. The reason is, that apart from EDM and laser ablation, milling and drilling are primarily applied for manufacturing molds. Examples of microfeature families are single-stage holes, free-form surfaces and plane surfaces. Fig. 18-6, for instance, depicts a set of demonstrator microparts which embody a realization of the microfeature ‘singlestage hole’. At the primary level of the catalog, the microfeatures are displayed together with their geometric and functional design data with the corresponding technical drawings. At the second level, geometric nominal values and tolerances and the nominal surface characteristics are juxtaposed on the actual values. During the comparison of these values, access to the process settings of the production process and the dimensional metrology is possible at the same time, since the allocation of target and actual values of the microfeatures are directly linked to relevant information on the specific production process and the applied measurement methods and strategies (see Fig. 18-7). This integration is essential in order to assure the reproducibility of both the production process settings and the chosen measurement strategy. In the context of quality assurance as a whole, microfeatures establish universal data objects recording quality results of all subprocesses involved. The knowledge is structured both design-specific and process-related. With the prototypically realized information front-end for the user, the microfeature catalog, the next steps are maintenance and improvement. Building on the constantly growing knowledge base, the adaptation and application of quality assurance methods will be carried out. Thereby, the already established regular meetings of quality control groups within the SFB 499, where specialists in design, production, analysis and quality assurance participate in order to facilitate information transfer and efficient cooperation, will be used for the primary application of adapted quality methods.

Fig. 18-7 Juxtaposition of nominal values and actual values of the microfeature ‘nozzle plate drilling’

18.4 Approaches to Quality Assurance and Dimensional Metrology in the SFB 499 511

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18.4.2 Dimensional Metrology in the SFB 499

Since dimensional metrology was not yet available as an adequate technology as such (see Section 18.3.2), one core objective of the subproject C4 was to choose an adequate technology and put it into operation. The chosen measurement machine is a high-precision coordinate measuring machine (CMM) equipped with the WFP (see Fig. 18-8) as an adequate commercially available measuring sensor for dimensional micrometrology (see Section 18.3.2). The basic precision of the machine is high, the probing elements are sufficiently small for most of the microstructures, the probing forces are negligible (< 0.1 lN) and the measurement of 2.5D structures is possible. Moreover, an image processing sensor with 10 ´ optics and a resolution of 0.05 lm together with an autofocus sensor are integrated. This multisensoric approach allows a flexible measurement of the microparts at hand. Each sensor is applied for the probing of those structural details for which it is most suitable. The dimensions of both molds and molded micromechanical parts of the demonstrator system are measured. Mainly, two sensors are applied for the actual measurement. On the one hand, information on planes, which are positioned parallel to the measurement table, is acquired using the autofocus senor. On the other hand, these results are complemented by tactile probing with the WFP. The image processing sensor is only employed for a first rough orientation of the measurement object before the actual measurement takes place, since optical sensors offer lower accuracy owing to unclear edge recognition and dependence on the optical properties of the material. This is even more relevant, since measurement processes have to be carried out automatically in order to prevent both the time-intensive application and the accuracy-decreasing influence of the user. Moreover, light parameters usually have to be adjusted manually from workpiece to workpiece.

Fig. 18-8 The high-precision measuring machine with the WFP, a schematic view of the opto-tactile sensor and a view of the sensor while probing a microgear

18.4 Approaches to Quality Assurance and Dimensional Metrology in the SFB 499

18.4.2.1 Measurement Results Regarding the manufacturing of molds, it is essential to carry out a 100% inspection for the dimensional, shape and position tolerances given in the technical drawing, since the geometry of the mold defines the geometry of the final parts. In particular, position tolerances are critical. For example, if a hole is not positioned correctly in the center of the mold of a gear, the resulting axis will not be positioned centrally in the molded part either. Currently, the accuracy of micromilling lies within ± 5 lm of the nominal value. Concerning the molding processes microPIM and microcasting, dimensional measurements are used on the one hand to inspect the dimensional accuracy of the final parts. On the other hand, information on molding accuracy and shrinkage is extracted in order to support the improvement of the molding processes. Consequently, not only final parts but also parts in their intermediate states are measured. Concerning the microPIM process, the mold, green parts and sintered parts have been measured. Owing to their fragility, brown parts have not yet been measured dimensionally. Even if it were possible to fix them for measurement, it would be impossible to remove a brown part without destroying it. With regard to microcasting, the mold, lost plastic patterns and final parts are measured. In Table 18-1, measurement results are shown exemplarily for both the microPIM and the microcasting processes. It should be mentioned that no statistically significant conclusions are possible, as the random sample is not sufficiently large. The reason is that the molding processes are in a constant train of improvement so that it is only possible to compare parts within one lot. To get an impression of the order of magnitude of dimensional accuracies, two examples are given. For the microcasting process different diameters of a mold and corresponding plastic patterns and final parts made of Stabilor G were measured. The molding accuracy comparing mold and plastic patterns was 0.4% and the shrinkage comparing plastic patterns and sintered parts was 0.01% for the given measured objects. For microPIM, a similar analysis was carried out for a mold and the corresponding green and sintered parts made of ZrO2. The molding accuracy comparing mold and green part was 0.1% in this particular case whereas the shrinkage comparing green part to sintered part was 22.7%.

Table 18-1 Examples for measurement results for the microPIM and microcasting processes

Process

Molding accuracy (%)

Shrinkage (%)

MicroPIM Microcasting

Mold – green part: 0.1 Mold – plastic pattern: 0.4

Green part – sintered part: 22.7 Plastic pattern – final part: 0.01

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18.4.2.2 Measurement Uncertainty As tolerances of micromechanical parts are in the micrometer or even sub-micrometer range, the quality of the measurement itself is a prevailing issue in particular for the opto-tactile microsensor, which is mainly applied. Accordingly, it is essential to specify the corresponding measurement uncertainty (abbreviated uncertainty), since this is the internationally standardized parameter assessing the quality of a measurement [30]. Only if the uncertainty of a measurement is known, the correct interpretation of measurement results is possible, e.g. when verifying compliance with given tolerances. Furthermore, different measurement results become comparable by means of this uniform measure due to given traceability. How to determine uncertainty is uniformly defined for all kinds of measurement in the Guide to the Expression of Uncertainty in Measurement (GUM) [30]. The core of the method is to quantify the effects of all relevant influence factors of a measurement on the measurement result. Since CMMs are universally applicable, each measurement is affected by a different set of influence factors. Hence the CMM uncertainty is task specific. Five different methods to derive this task-specific CMM uncertainty have been developed [31]. As the core task of the subproject C4 is not the further development of micrometrology, but rather its application, the seemingly most feasible and user-friendly method for CMM uncertainty estimation was selected. The method chosen for adaptation to the micrometer regime was the so-called experimental method according to ISO/TS 15530-3 and DIN 32881-3 [32].

The Experimental Uncertainty Estimation The estimation of the task-specific CMM uncertainty with the experimental approach according to the standards ISO/TS 15530-3 and DIN 32881-3 [32] is based on the repeated measurement of one or more calibrated objects which resemble the actually measured objects. These repeated measurements are executed under the same conditions and in the same way as the actual measurement for which the uncertainty estimation is carried out. For the final estimation of the expanded uncertainty U, three different uncertainty sources, uC, uP and uW (see Fig. 18-9), are taken into account and the expanded uncertainty U is calculated according to the formula given in Fig. 18-9.

Fig. 18-9 Calculation of the expanded uncertainty U according to DIN 32881-3

18.4 Approaches to Quality Assurance and Dimensional Metrology in the SFB 499

Fig. 18-10 Experimental estimation of uncertainty

The components of the expanded uncertainty for the measurement of one specific measured value are derived as follows (see Fig. 18-10). The starting point is the selection of one or more reference objects and the corresponding measured value taking into account the following two constraints. On the one hand, high similarity requirements between the actual objects and the reference object(s) in terms of geometry, material and production process have to be met to ensure adequate results of the uncertainty estimation [32]. Optimally, one of the actually measured objects is used for calibration. On the other hand, the form error of the reference object(s) has to be significantly lower than the value of the measured value for calibration. After having selected the reference object(s), the calibration method has to be chosen. Since the uncertainty of calibration equals uC and this directly contributes to the expanded uncertainty U (see Fig. 18-9), the uncertainty of calibration should be negligibly small compared with the expected uncertainty of the given measurement device. Once the reference objects have been calibrated, a minimum of 20 measurements [32] are carried out under the same conditions and with the same measurement strategy as the actual measurement. The statistical standard deviation of the results obtained equals the standard uncertainty of the measurement process uP, and the systematic error E is equal to the difference of the mean of the results and the calibration value. If the systematic error E is corrected, it does not contribute to the uncertainty calculation. Since workpiece features such as material, form, shape and surface influence measurements, the parameter uW covers the influence of all relevant variations of workpiece features due to variations of the production process and the material. Unless the reference objects resemble the actually measured objects within negligible variations, uW has to be quantified. This implies deriving the interrelation between the deviation of each influencing factor and the deviation of the measurement result either statistically or based on experience. Finally, the interdependences of the influencing factors have to be determined to calculate the resulting combined standard uncertainty uW.

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Application of the Experimental Uncertainty Estimation Since the uncertainty of coordinate measurements is task specific, the uncertainty actually already differs if only one part of the measurement strategy is altered or the measurement task is slightly different. Consequently, to carry out an uncertainty estimation for each measurement task would require extensive effort. Therefore, the selected measured values are the diameter of a hole of 120 lm and the distance of two planes of 200 lm, since their measurement is a typical task in the dimensional 2.5D micromeasurement of the considered microparts. However, it was not possible to calibrate reference objects from the given production owing to an insufficient form accuracy. As a consequence, at least for the hole, a substitute in microdimensions, which was provided by the Physikalisch-Technische Bundesanstalt (PTB), was found. The substitutional reference object was a 0.4 lm thin metal disk with a hole of 200 lm diameter and a form deviation of < 0.4 lm. Even though this object offers a sufficient form error for calibration, the obvious drawback is the lack of similarity to the actually measured objects in terms of both material and production process. Regarding the calibration method, a problem which is typical for micrometrology occurred. Since particularly with regard to 3D micrometrology the possibilities of measurement are limited, given a certain measurement device such as the opto-tactile sensor it is difficult to find a 2.5D calibration method with a comparably lower uncertainty uC. In this case, the calibration had to be carried out with the same sensor, i.e. the opto-tactile sensor. As an undesirable effect, the uncertainty of the opto-tactile measurement contributes twice to U, once in uC and once in uP (see Fig. 18-9). The uncertainty of calibration uC equals 0.5 lm. In the next step, the measurement of the calibrated hole was carried out 20 times according to the actual measurement strategy, at the same time varying the position of the measurement object systematically over the CMM high-precision measuring table. The resulting uP equaled 0.31 lm. Regarding the specification of the parameter uW for an opto-tactile measurement, an examination of the influence factors form, surface roughness, porosity, adhesion, elasticity and the coefficient of expansion has to be carried out. It is not possible to model the dependence of these influence factors on the measurement result without extensive experience, knowledge and effort. As only an equation for the estimation of the uncertainty contribution resulting from variations of the coefficient of expansion is given in DIN 32881-3, only rough estimations of uW were possible. Overall, the result of applying the experimental approach for uncertainty estimation primarily was an improved understanding of the uncertainty influences. As described above, two obstacles regarding calibration arose, i.e. selecting an adequate reference object for calibration and finding an available calibration method with a sufficiently low uncertainty. Additionally, the specification of the parameter uW is complex. Therefore, the estimation only gives a first clue to the quantified uncertainty of WFP measurements. Based on the information that uC equals 0.5 lm for the same sensor and the same machine, the expanded uncertainty U of measurements within the subproject C4 is at least 1 lm. Consequently, for tolerances of only a few micrometers, a decisive inspection based

18.6 Acknowledgment

on dimensional measurement results is not feasible in most cases. As a result, dimensional measurements are carried out in an internally standardized way so that measurement results are comparable among each other. Moreover, they will be complemented by functional tests within the SFB 499.

18.5 Conclusions and Outlook

In the field of quality management, quality assurance as a both preventive and operative production-accompanying process is of particular importance in the production of micromechanical parts. Owing to the lack of experience and know-how and challenges regarding high-precision 3D-applicable micrometrology, the main objectives are the development of concepts and structures for a microspecific information–technological support of all subprocesses and the establishment of a robust dimensional measurement process. As a consequence, the major research within the subproject C4 has been concentrated on these two topics. Microfeatures and the corresponding information tool, the microfeature catalog, have been developed to permit structured, process-related and microspecific information storage and provision. Regarding the dimensional measurement, measurements have been optimized based on the experience gained by applying the experimental method. For the future, the adaptation and application of quality assurance methods within the production of micromechanical parts are of paramount importance in order to adhere fully to the specifics of microproduction and thus guarantee adequate support. Moreover, economically efficient process-spanning quality control loops have to be established to guarantee holistic quality support along the whole process chain.

18.6 Acknowledgment

This research was financially supported by the German Research Foundation (DFG) within the Collaborative Research Centre (SFB) 499 ‘Design, Production and Quality Assurance of Primary Shaped Microcomponents Manufactured in Ceramic and Metallic Materials’.

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18.7 References 1 T. Pfeifer, Qualitätsmanagement, 3rd edn; 2

3

4

5

6

7

8

Munich: Carl Hanser, 2001. T. Pfeifer, R. Freudenberg, G. Dussler, B. Bröcher, Quality control and process observation for the microassembly process. Meas. J. Int. Meas. Confed. 2001, 30(1), 1–18. P. Knauth, P. Feitner, A. Geissel, Qualitätsmanagement in der mikrotechnischen Produktion. In: Projektbericht Fertigungstechnik für passive optische Komponenten, Ergebnisbericht des Verbundprojektes FEPAKOM, Projektträgerschaft Fertigungstechnik und Qualitätssicherung (PFT); Wissenschaftliche Berichte FZKA-RFT 185, Karlsruhe: FZK Karlsruhe, 1998, pp. 113–148. D. Maas, R. Ehners, P. J. Jakobs, B. Karl, U. Köhler, C. Lienert, V. Saile, J. Schulz, Fertigung von Mikrokomponenten nach DIN EN ISO 9001 an einem Forschungsinstitut. In: 4. Statuskolloquium des Programms Mikrosystemtechnik 30./31. März 2000: Karlsruhe: Institut für Mikrostrukturtechnik, FZK Karlsruhe, 2000, pp. 75– 79. S. Büttgenbach, ‘Rechnergestützte Entwurfsumgebung’ im Rahmen des SFB 516; TP A3, Braunschweig: Universität Braunschweig, Institut für Mikrotechnik (IMT), 2001. D. Zielke, R. Lieske, J. Will, Automatic transfer from bulk-silicon technology simulation into the FEM-environment. In: Transducers ’01/Eurosensors XV, 11th International Conference on Solid-State Sensors and Actuators, Munich, 11–14 June, 2001, Digest of Technical Papers; Vol. 1, 2001, pp. 272–275. T. Pfeifer, G. Dussler, S. Driessen, Confocal inspection of microscopic parts and structures. In: Proceedings of the 3rd International EUSPEN Conference, Eindhoven, 27–29 May 2002; 2002, pp. 505–508. T. Pfeifer, G. Dussler, Process observation for the assembly of hybrid microsystems. In: Proceedings for the 13th International Symposium on Micromechatronics and Human Science. Nagoya, 21–23 October 2002; 2002, pp. 117–123.

9 A. Weckenmann, R. Ernst, Anforderun-

10

11

12

13

14

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16 17

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20

gen an die zukünftige Mikro- und Nanomesstechnik. In: Technisches Messen 2000, 7–8, 334–342. Institut für Fertigungstechnik und Materialforschung (IFAM). Rasterelektronenmikroskopie. http//www.ifam.fhg.de/ index.php?seite=/2801/zentren/ materialografie/raster/; accessed 2 September 2004. Institut für Festkörperphysik Rastersondenmikroskopie. http//www.nano.unihannover.de/proseminar04/SXM.pdf; accessed 2 September 2004. A. Weckenmann, T. Estler, G. Peggs, D. McMurtry, Probing systems in dimensional metrology. Ann. CIRP 2004, 53(2), 1–28. M. Andräs, Zwei in einem. Ein neuer Mikrotaster für Koordinatenmessgeräte kombiniert taktile Antastung mit Bildverarbeitung. Qualität Zuverlässigkeit 1998, 43(6), 740. R. Christoph, H. J. Neumann, Multisensor-Koordinatenmesstechnik. Die Bibliothek der Technik, Vol. 248, 1st edn.; Neuss: Landsberg/Lech Verlag Moderne Industrie, 2003. Mitutoyo. http//www.mitutoyo.com.tw/ newsite/image/UMAP%20Vision.pdf; accessed 15 June 2004. Product Information. Giessen: Werth Messtechnik, 2004. Mitutoyo UMAP Vision System. Mitutoyo. http//www.mitutoyo.co.jp/eng/news/ newsfile/02data/02_13.html; accessed 27May 2004. S. Cao, U. Brand, B. T. Kleine, W. Hoffmann, H. Schwenke, S. Bütefisch, S. Büttgenbach, Recent developments in dimensional metrology for microsystem components. Microsyst. Technol. 2002, 8(3–6), 3–6. K. Enami, M. Hiraki, K. Takamasu, Nano-probe using optical sensing. In: XVI IMEKO World Congress – IMEKO, Vienna; 2000. K. Takamasu, C.-C. Kuo, A. Suzuki, M. Hiraki, R. Furutani, S. Ozono, Development of a pneumatic ball probe (1st re-

18.7 References

21

22

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24

25 26

port) – basic construction. Int. J. Jpn. Soc. Prec. Engl. 1999, 33(3), 233–238. U. Brand, S. Büttgenbach, V. Nesterov, P. Pornnoppadol, R. Wilke, 3D-Mikrotaster. In: Metrologie in der Mikro- und Nanotechnik; VDI-Berichte 1669, Düsseldorf: VDI Verlag, 2003, pp. 117–126. PTB Opto-taktiler Taster zur Messung kleiner Bohrungen und 3D Mikrostrukturen. http//www.ptb.de/de/org/5/53/ f_vorhaben/optotakiler_taster.html; accessed 28 January 2004. PTB Projekt 5.101 Mikrosystem-Messtechnik. http//www.ptb.de/de/org/5/51/ 5101/seite_3d.htm; accessed 28 January 2004. A. R. Grayer, A computer link between design and manufacture. PhD Thesis; University of Cambridge, 1976. VDI Directive 2218 Feature–Technologie; Berlin: Springer, 1998. K. Hain, Automatische Gewinnung von Merkmalen und Klassifizierungseigenschaften für Produkte auf Basis eines integrierten Produktmodells. In: Forschungsberichte aus dem RPK (Karlsruhe), Vol. 97; Shaker Verlag, 1997.

27 D. Spath, J. Elsner, Quality assurance of

28 29

30 31

32

micromechanical components. In: Proceedings of European Society for Precision Engineering and Nanotechnology, Eindhoven, 26–29 May 2002; 2002, pp. 403–407. N. P. Suh, The Principles of Design; New York: Oxford University Press, 1990. F. Fleissner, Prozessorientierte Prüfplanung auf Basis von Bearbeitungsobjekten für die Kleinserienfertigung am Beispiel der Bohr- und Fräsbearbeitung. Dissertation; Institute for Machine Tools and Production Science, University of Karlsruhe, 1998. ISO Guide to the Expression of Uncertainty in Measurement; ISO, 1995. R. G. Wilhelm, R. Hocken, H. Schwenke, Task specific uncertainty in coordinate measurement. Ann. CIRP 2001, 50(2), 553–563. Verfahren zur Bestimmung der Messunsicherheit von Koordinaten-messgeräten (KMG) – Teil 3. Unsicherheitsermittlung mit kalibrierten Werkstücken. DIN 32881-3; 2000, Berlin: Beuth Verlag.

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VI Properties of Materials and Microcomponents

Advanced Micro and Nanosystems Vol. 4. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31493-8

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19 Analysis of Microstructure, Surface Topography and Mechanical Properties of Microcast Specimens Made of the Dental Gold Alloy Stabilor G B. Kasanická, M. Auhorn, V. Schulze, T. Beck, D. Löhe, Institute of Material Science and Engineering I, University of Karlsruhe (TH), Germany Abstract Examinations of specimens made of the gold-based dental alloy Stabilor G produced by microcasting are presented. Two different casting techniques were applied: vacuum pressure casting and centrifugal casting. In order to study the influence of the parameters of the casting process such as the cooling conditions and the filling pressure, the grain size, the surface topography, the roughness and the mechanical properties of microcast specimens were varied. These results allow one to understand correlations between process parameters, the characteristics of the microspecimens and plasticity effects at small scale such as hardening mechanisms and anisotropy, which are useful for improving the production process.

Keywords microcasting; gold-based dental alloy; microstructure; surface topography; mechanical properties

19.1 19.2 19.2.1 19.2.2 19.3 19.3.1 19.3.2 19.3.2.1 19.3.2.2 19.3.3 19.4 19.5 19.5.1

Introduction 524 Theoretical Background

524

Relevant Hardening Mechanisms 524 Deformation Behavior of Single and Polycrystals

526

Analyzed Material States 526 Analyzed Material 526 Production Process 526

Vacuum Pressure Casting (VPC) Centrifugal Casting (CC) 527 Cooling Conditions 528

527

Characterization Techniques 528 Results 531

Microstructure of Microsticks after Vacuum Pressure Casting

Advanced Micro and Nanosystems Vol. 4. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31493-8

531

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19 Analysis of Microstructure, Surface Topography and Mechanical Properties

19.5.2 19.5.2.1 19.5.2.2 19.5.2.3 19.5.2.4 19.5.3 19.5.3.1 19.5.3.2 19.6 19.7

Microstructure and Mechanical Properties of Microtensile Specimens 533 Microtensile Specimens Produced by VPC 533 Microtensile Specimens Produced by CC 538 Comparison of Microtensile Specimens Produced by CC and by VPC 543 Microtensile Specimens Produced by CC with Additional Annealing at 8008C for 5 h 544 Microstructure and Mechanical Properties of Microbending Specimens 547 Quasi-static Micro-three-point Bending 547 Cyclic Micro-three-point Bending 549 Conclusion and Outlook References 552

552

19.1 Introduction

Microcasting is a very suitable process for the shaping of microcomponents made of metal alloys. The advantage of microcasting compared with, e.g., the LIGA technique is the possibility of producing real 3D microcomponents and to apply distinctly more materials to the production process [1]. Owing to its very good castability and the broad knowledge about its behavior in small-scale applications such as in dental practice, the gold alloy Stabilor G (Au58Ag23Cu12Pd5) was one of the first choices for microcasting investigations at the Institut für Materialforschung III of the Forschungszentrum Karlsruhe (see Chapter 13). In the investigations presented in this chapter, microstructure and surface topography of microcast specimens made of the gold-based dental alloy Stabilor G will be correlated with the behavior at quasi-static tensile loading. Additionally, a short overview of mechanical testing devices for microspecimens will be given. Finally, the results will be discussed and evaluated based on the effects of different hardening mechanisms and the effects of elastic and plastic anisotropy caused by the small number of grains.

19.2 Theoretical Background 19.2.1 Relevant Hardening Mechanisms

It is known that an age-hardening phenomenon occurs in several types of gold dental alloys depending on the contents of silver and copper. This phenomenon can be utilized deliberately to harden these alloys and improve their mechanical

19.2 Theoretical Background

properties [2, 3]. There are three major hardening mechanisms: precipitation hardening, hardening by ordering and grain boundary hardening. Precipitation hardening can be explained in a simplified way by means of an example. If an AgCu alloy containing sufficient copper is slowly cooled (from 700 8C to room temperature), the surplus of copper precipitates from the silverrich matrix as a second phase – a copper-rich solid solution – owing to the decreased solubility for copper. If the alloy is rapidly cooled by water quenching, such precipitation is suppressed and a supersaturated solution with copper results. Subsequently, heat treatment at about 200–400 8C permits the precipitation of the copper-rich second phase in a very finely dispersed form [3]. This increases the hardness significantly because the precipitates are obstacles for the dislocation movement necessary for plastic deformation. The mechanism of overcoming the obstacles depends on the coherence of the precipitates. If they are small and coherent, the dislocations need to cut them. If they are large or incoherent, they will be overcome by Orowan looping. Hardening by ordering is a special mechanism occurring in gold–copper alloys and in yellow AuAgCu alloys where the copper content is relatively high. In the gold–copper system, a homogeneous solid solution exists at all concentrations above *410 8C. The gold and copper atoms are randomly distributed in the alloy crystal lattice. However, below this temperature, several ordered superstructures will be formed. The copper and gold atoms occupy fixed positions in the crystal lattice and the alloy is harder and less ductile [3]. It has been shown that the formation of a CuAu I superlattice is essential to the age-hardening process and that precipitation plays a subordinate role [2]. The mechanism leading to the hardening effect is the necessity for coordinated movement of two dislocations within the ordered precipitate because the movement of a single dislocation will lead to an anti-phase boundary. Grain boundary hardening is usually of less relevance in gold alloys than the above-described hardening mechanisms. As grain boundaries are obstacles for dislocation movement which cannot be overcome and lead to dislocation pileups that may induce dislocation movement in neighboring grains, the hardness and yield stress of the material typically increase with decreasing grain size, a phenomenon which can be described by the Hall–Petch relation [4, 5]: ry ˆ r0 ‡ kd

1=2

…1†

where ry is the yield stress, r0 is the intrinsic yield stress which represents the result of all the other hardening mechanisms, k is a material-dependent constant and d is the average grain size. Therefore, to achieve a higher hardness and yield stress for gold-based alloys, the grains are refined mainly by alloying small amounts of iridium segregating at the grain boundaries [6].

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19.2.2 Deformation Behavior of Single and Polycrystals

The deformation of crystalline materials can be classified roughly in single and polycrystalline material behavior. Single crystals show an anisotropic material behavior that is heavily influenced by the relation of crystal orientation to loading direction, which means that the stiffness of a single crystal usually differs with loading direction. This can be illustrated very well with the ratio of maximum to minimum Young’s modulus of, e.g., gold: Eh111i/Eh100i = 114 GPa/ 42 GPa = 2.71 [6]. In contrast, polycrystalline materials that have randomly distributed grain orientations show a material behavior that is independent of the loading direction; the material behavior is then called quasi-isotropic. Therefore, polycrystalline materials are characterized with only a single Young’s modulus, e.g. for gold E = 78 GPa [7]. Placed between single and polycrystals there are bicrystals and compounds of several crystals (grains) whose material behavior is heavily influenced by the anisotropy of the individual grains. Owing to the small number of grains, there is no statistical averaging of the grain orientations. By loading single crystals or multi-grain compounds in tension, the individual grains start to rotate in such a way that the slip direction approaches the tensile direction [8]. This results, in the case of a grain compound, in different stress and strain levels in the grain compound and/or in individual grains depending on the local grain orientation and the interaction with neighboring grains [9, 10]. Investigations on the influence of different grain orientations on the local stress distribution in relation to different numbers of grains were performed by finite element simulations (FEM) and are presented in Chapter 2.

19.3 Analyzed Material States 19.3.1 Analyzed Material

The examined material was a commercially available gold-based dental alloy, Stabilor G, which has the chemical composition presented in Table 19-1. The melting interval of this alloy is 860–940 8C. 19.3.2 Production Process

Microsticks and microbending specimens (both 250 ´ 250 ´ 1200 lm) were produced by vacuum pressure casting and microtensile specimens (130 ´ 260 ´ 4500 lm, gauge length 780 lm) by vacuum pressure casting and by centrifugal casting using facilities located at the Institut für Materialforschung III of the Forschungszentrum Karlsruhe, Germany (see Chapter 13). In both

19.3 Analyzed Material States Table 19-1 Chemical composition of Stabilor G in wt.% and atom% (values given by manufacturer)

Element

Wt.% Atom%

Au

Ag

Cu

Pd

Zn

Pt

Ir

58.0 38.4

23.3 28.2

12.0 24.6

5.5 6.7

1.0 2.0

0.1 0.1

0.1 0.1

casting methods, models made of PMMA are mounted on a gate and feeding system made of wax. This assembly is then embedded in Deguvest, a commercially available phosphate-bonded slip used in the manufacture of ceramic molds for the dental casting of precious metal alloys. Additionally, plaster was used as mold material for a single batch. After drying, the ceramic is sintered to obtain a ceramic mold. During the sintering process, the plastic melts and is pyrolyzed. The metal microparts are then produced by filling the mold with the metal melt [1, 11].

19.3.2.1 Vacuum Pressure Casting (VPC) This casting process entails first fitting the preheated mold directly above a crucible which contains the metal melt in the casting equipment. The system is then evacuated and the mobile casting chamber rotated by 1808 so that the preheated mold is inverted and the metal melt flows into the mold by gravity. The casting chamber is then pressurized at 3–4 bar using a gas atmosphere in order to achieve complete form filling. The entire casting process takes 40 s. Thereafter, the filled mold is removed from the casting equipment and allowed to cool in air. Then the cast part is cleaned from the ceramic mold material with hydrofluoric acid. This is followed by separating the microspecimens from the gate and feeding system (see Chapter 13).

19.3.2.2 Centrifugal Casting (CC) In the centrifugal casting process, the preheated mold is first fitted into the casting chamber. The equipment is then evacuated and flushed with argon. This preparation process takes 3 min. An arc is ignited to heat the metal melt, which then flows through a runner in the mold. Both the runner and the mold rotate at 3000 min–1. Therefore, the form filling is performed in a pressure range of 20–25 bar. Then the casting process is stopped, the mold is removed from the casting equipment and kept in air to cool. The subsequent cleaning and singling process is the same as in vacuum pressure casting (see Chapter 13).

527

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19 Analysis of Microstructure, Surface Topography and Mechanical Properties

19.3.3 Cooling Conditions

The parameters of the casting process such as the preheating temperature of the mold and the filling pressure determine the grain size, the surface topography, the roughness and the mechanical properties of the specimens. In order to analyze the influence of the preheating temperature on the microstructure and the mechanical properties of the microspecimens, different preheating temperatures of the mold, Tm, and of the runner, Tr, were applied. In order to analyze the formation of the CuAu I superstructure, microtensile specimens of a single batch were additionally annealed in the vacuum furnace and subsequently cooled in the furnace from the annealing temperature to room temperature in 2 h. An overview of cooling and annealing conditions is presented in Table 19-2.

19.4 Characterization Techniques

The microstructure of the microsticks and microtensile and microbending specimens produced by different cooling conditions was examined using a light microscope (Leica Aristomet) and a transmission electron microscope (Zeiss EM 912 Omega). Prior to the microstructural investigation using an optical microscope, the microspecimens and their gates were chemically etched using a mixture of 100 ml of H2O, 100 ml of H2O2 and 32 g of FeCl3 as etchant. The grain

Table 19-2 The type of microspecimen, the casting technique, the cooling conditions and the material type of the mold for all batches investigated

Batch

Type of microspecimen

Casting technique

Cooling conditions

Material type of mold

A B C D E F G H I J

Microstick Microstick Microstick Microstick Tensile Tensile Tensile Tensile Tensile Tensile

VPC VPC VPC VPC VPC VPC CC CC CC CC

Deguvest Deguvest Deguvest Deguvest Deguvest Plaster Deguvest Deguvest Deguvest Deguvest

K

Bending

VPC

Tm = 100 8C Tm = 400 8C Tm = 700 8C Tm = 1000 8C Tm = 1000 8C Tm = 700 8C Tm = 700 8C Tm = 700 8C, Tr = 700 8C Tm = 1000 8C Tm = 1000 8C, annealed at 800 8C, 5 h Tm = 700 8C, water quenched, annealed at 400 8C, 20–30 min

Deguvest

19.4 Characterization Techniques

size was determined by applying the intercepted-segment method. The porosity was measured by using the binary method, which entails the optical separation of a gray scale spectrum using computer-aided algorithms. For the analysis of microstructure using transmission electron microscopy, the microspecimens were inserted in a disc which was cut from a dagged cylinder (Fig. 19-1) and then fixed with epoxy resin. Subsequently the microspecimens were conventionally dimpled by mechanical grinding and then thinned to electron transparency using Ar+ bombardment [12]. Additionally, the surface topographies of the mechanically investigated microspecimens were investigated using a confocal white light microscope (Nanofocus lSurf) with a vertical resolution of 0.1 lm and a lateral resolution of 1.5 lm [13]. From the surface topographies the relating peak-to-valley heights Rt were measured. However, the Rt values could only be determined in the longitudinal specimen direction according to DIN EN ISO 4288 [14], which recommends a minimum measuring length of 0.48 mm. Additionally, this standard suggests a cut-off wavelength of 250 lm in order to distinguish the roughness from the waviness. Therefore, it cannot be completely excluded that some of the widespread surface defects of the specimens are recognized as parts of the waviness profile, which leads to slightly underestimated Rt values in some cases. As already mentioned, the experimental determination of the mechanical properties of microsized specimens is of great relevance in order to optimize production processes and to dimension microparts. Therefore, bending and tensile tests are adapted for the microrange. Examples of micromechanical tests known from the literature are tensile tests on Si (specimens produced by wet etching) [15], Ni and Ni–Fe [15, 16] with gauge dimensions between 50 ´ 70 ´ 320 and 100 ´ 100 ´ 1000 lm and absolute specimen lengths of 10 or 16 mm. In [16], for the microtensile tests a conventional macroscopic tensile testing machine with an additionally installed load frame is used. The axial strain is determined as an integral value by measuring the distance of two specimen shoulders with a laser extensometer. The functional principle of the testing device used in [15] is a combination of a micrometer positioning device and

Fig. 19-1 The fixed disc with a microbending specimen (left) and a view of the microbending specimen prepared for transmission electron microscopic investigations (right)

529

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19 Analysis of Microstructure, Surface Topography and Mechanical Properties

a piezo column; the locally resolved axial strain is quantified by analyzing light optical microscope images of the testing region. Additionally, this device is also used for microbending tests by applying the same specimen design. Another example of microbending tests is the testing device described in [17]: a magnetostrictive actuator is used to test Ni–P microcantilever beam specimens which have the dimensions 10 ´ 12 ´ 50 lm (specimens produced by focused ion beam machining). Additionally, new types of tests are established in order to obtain information about local mechanical properties by nanoindentation [18–20]. More detailed overviews of micromechanical investigations are given in [21–24]. The mechanical properties of the microspecimens presented here were determined using a self-designed universal microtesting machine [25, 26]. The machine has an electrodynamic actuator and allows tests with both quasi-static and cyclic loadings. It can be equipped with microtensile clamps or with a microthree-point bending device, both produced by microelectrodischarge machining. The lower bending supports have a distance of l = 0.8 mm and a radius of 50 lm. The upper support has a radius of 100 lm. Further descriptions of the microbending device are given in [27] and in Chapter 20. The maximum force of the testing machine is *50 N. The force is measured with a commercially available 100 N load cell and the displacement by a capacitive path-measuring system. Before starting the mechanical tests, the dimensions of the cross sections of the individual specimens have to be measured using the already mentioned confocal white-light microscope. All quasi-static experiments are performed in path control by raising the displacement linearly at a velocity of 2 lm/s until fracture occurs. Closed loop control of the strain is not possible using samples with such small dimensions. However, strain fields at the specimen surfaces are determined after the tests by analyzing optical microscope images, which are recorded continuously during the experiments using commercial gray-scale correlation software. Subsequently, the software compares the images, determines local displacements of optical contrasts on the specimen surface and transforms the measured displacements into strain values [28]. Cyclic bending tests are carried out in force control and start with linearly raising the force with time up to a mean value, which is kept constant. Subsequently, a sinusoidal cyclic loading is superimposed with a frequency of 10 Hz. The nominal tensile stress rn is calculated as usual by relating the force F to the initial cross section A0 according to the equation rn ˆ

F A0

…2†

By this procedure, the yield strength ry and the ultimate tensile strength rm are determined. The Young’s modulus Ei is calculated by determining an integral strain value eel, longitudinal, i from the locally resolved strain field belonging to a sufficiently low loading so that pure elastic deformation can be assumed and by relating this integral strain value to the corresponding nominal stress rn,el,i:

19.5 Results

Ei ˆ

rn;el;i eel;longitudinal;i

…3†

From the evaluation of several images during the purely elastic loading, a mean value E of the Young’s moduli Ei is calculated. Poisson’s ratio is determined by relating the integral values of the total strains in the purely elastic regime in lateral and longitudinal direction: vi ˆ

eel;lateral;i eel;longitudinal;i

…4†

Corresponding to Young’s modulus, a mean of the values at several loading steps in the elastic regime is calculated. The elongation to fracture A is derived by determining the integral value of the total strain from the last image before fracture occurs et,f and by subtracting the fracture stress rf divided by the mean value of the Young’s modulus E: A ˆ et;f

rf E

…5†

The bending stresses are calculated for both elastic and elasto-plastic material behavior purely elastically for the margin of a specimen at the point of highest load as follows: 3 F
l rb ˆ
2 b
h2

…6†

where F is the centric force applied to the three-point bending specimen, l is the distance of the bending supports, i.e. 0.8 mm, and b and h are the cross-sectional dimensions breadth and height.

19.5 Results 19.5.1 Microstructure of Microsticks after Vacuum Pressure Casting

The temperature of the preheated mold, the grain size and porosity of all batches of the microsticks are presented in Table 19-3. In Fig. 19-2, upper left, the microstructure of a gate (Tm = 100 8C) is presented (batch A). Because of the low temperature of the ceramic mold, it was not possible to obtain a completely filled microstick from this batch for microstructure analysis. As a consequence of the high cooling rate of the metal melt, the gate shows a characteristic fine dendritic cast microstructure. Fig. 19-2, upper right, shows the microstructure

531

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19 Analysis of Microstructure, Surface Topography and Mechanical Properties Table 19-3 Temperatures of the preheated mold, grain size and porosity of the batches of the investigated vacuum pressure cast microsticks

Batch

Temperature of preheated mold (8C)

Grain size (lm)

Porosity (%)

A B C D

100 400 700 1000

– – 17.4 42.8

0.5 0.1 0.2 0.3

Fig. 19-2 Microstructure of the gate (A) and of the microsticks (B–D) produced by VPC using a preheated mold at Tm = 100 (A), 400 (B), 700 (C) and 1000 8C (D) [29]

of the microstick which was cast using a preheated mold at Tm = 400 8C (batch B). In comparison with Fig. 19-2, upper left, this figure shows coarser dendrites. The dark areas represent segregation zones, which is in agreement with [30]. Because of the higher temperature of the preheated mold and thereby the lower rate of solidification, the weaker secondary and tertiary arms of dendrites reduce for the benefit of the primary dendrites, which coarsen. Owing to the dendritic structure of the microsticks of batches A and B, the grain size could not be determined. In Fig. 19-2, lower left, the microstructure of the microstick (Tm = 700 8C) is presented (batch C). Because of the higher temperature of the

19.5 Results

preheated mold, the constitutional supercooling of the material is reduced. The predominantly homogeneous microstructure shows only very small areas of segregation and a grain size of 17.4 lm. In addition to the free grain boundaries, boundaries decorated with a secondary phase were also observed. In additional studies using transmission electron microscopy (TEM), it was found that this phase contains subgrains, twins and dislocations. Fig. 19-2, lower right, shows the microstructure of the microstick which was cast at Tm = 1000 8C (batch D). The microstructure shows only a few grain boundaries with the secondary phase and a distinctive coarsening of the grains. The grain size was determined as 40.8 lm and no segregation was observed inside the grains. A typical arrangement of the grains in a cross section of macrocomponents studied in [30] was composed of three areas: a peripheral zone with very small grains, an intermediate crystallization zone with columnar dendrites and a core zone with globular–dendritic coarsened grains. In the microspecimens, this typical macroscopic cast microstructure was not observed because of the much smaller thermal gradient in the 200 ´ 200 lm cross section of the microsticks. All four batches showed low porosities between 0.1% (batch B) and 0.5% (batch A). 19.5.2 Microstructure and Mechanical Properties of Microtensile Specimens

The casting technique, the cooling conditions, the material type of the mold, the grain size and the porosity of all the batches of the microtensile specimens are presented in Table 19-4. The results of the mechanical tests on the microtensile specimens are presented in the following paragraphs in Tables 19-5–19-10.

19.5.2.1 Microtensile Specimens Produced by VPC The microtensile specimens of batch E (Tm = 1000 8C) show a very homogeneous microstructure (Fig. 19-3, left). The grain size was determined as 64.0 lm. BeTable 19-4 Casting technique, cooling conditions, material type of the mold, grain size and porosity of batches E–J of vacuum pressure and centrifugal cast microtensile specimens

Batch Casting Cooling conditions technique

E F G H I J

VPC VPC CC CC CC CC

Tm = 1000 8C Tm = 700 8C Tm = 700 8C Tm = 700 8C, Tr = 700 8C Tm = 1000 8C Tm = 1000 8C; annealed at 800 8C, 5 h

Material type of mold

Grain size Porosity (lm) (%)

Peak-to-valley roughness (lm)

Deguvest Plaster Deguvest Deguvest Deguvest Deguvest

64.0 35.1 10.7 11.2 10.2 35.9

4.3 1.7 8.7 7.5 7.9 6.4

0.2 0.2 0.0 0.1 0.7 0.0

533

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19 Analysis of Microstructure, Surface Topography and Mechanical Properties

Fig. 19-3 Microstructure of the microtensile specimens produced by VPC at Tm = 1000 8C using a Deguvest mold (E) and at Tm = 700 8C using a plaster mold (F)

cause of the larger amount of the metal melt (m = 16 g), the specimens show a larger grain size than the respective microsticks produced by VPC using a preheated mold at the same temperature and an amount of metal melt of m = 10 g (batch D). Fig. 19-3, right, shows the microstructure of the microtensile specimens of batch F (Tm = 700 8C). Because of the lower temperature of the preheated mold, the grain size is reduced to 35.1 lm. In the microstructures of both batches, agglomerates of Ir particles were observed. The presence of these agglomerates can have falsified the result of the porosity measurement because of the difficulty in keeping the pores and the Ir particles apart. By image analysis, a porosity of 0.2% for both batches was obtained. The peak-to-valley roughness of the microtensile specimens of batch E was determined as 4.3 lm. The surface topography of these microspecimens is shown in Fig. 19-4, left. After VPC using a

Fig. 19-4 Surface topography of the microtensile specimens produced by VPC at Tm = 1000 8C using a Deguvest mold (batch E, left) and at Tm = 700 8C using a plaster mold (batch F, right)

19.5 Results

plaster mold specially developed for casting an aluminum bronze alloy, the microtensile specimens of batch F show very smooth surfaces (Fig. 19-4, right) with a peak-to-valley roughness of 1.7 lm. The stress–strain curves for five specimens of batch E are shown in Fig. 19-5. The related mechanical properties are given in Table 19-5. The slight ups and downs at the end of some curves are due to changes in the specimen compliance and its effect on the closed loop control and are neglected in further discussion. The stress–strain curves and the mechanical properties show a relatively large scatter of the material behavior of batch E. This scatter of all mechanical parameters, especially of the ry and E values, reveals that the deformation behavior is influenced by anisotropy. This can be additionally illustrated by analyzing the strain fields with regard to the local strain development. Therefore, in Fig. 19-6 two gray-scale correlation images are given for the specimen of batch E which showed the highest ultimate tensile strength of rm = 658 MPa. Fig. 19-6 a, which corresponds to the nominal stress rn = 377 MPa close to the yield strength ry = 375 MPa, reveals a very inhomogeneous strain distribution with local axial

Fig. 19-5 Stress–strain curves for batch E (VPC, Tm = 1000 8C)

Table 19-5 Mechanical properties of batch E (VPC, Tm = 1000 8C)

ry (MPa)

m

E (GPa)

Min.

Max.

Mean

Min.

Max.

Mean

Min.

Max.

Mean

307

388

354

59

99

79

0.24

0.49

0.35

rf (MPa)

rm (MPa)

A (%)

Min.

Max.

Mean

Min.

Max.

Mean

Min.

Max.

Mean

504

658

594

451

616

527

5.37

11.51

8.64

535

536

19 Analysis of Microstructure, Surface Topography and Mechanical Properties

Fig. 19-6 Surface image and optically resolved gray-scale correlation image (strain field) of the batch E specimen with rm = 658 MPa for rn = (a) 377 and (b) 607 MPa

strain values between –0.4 and 1.1%. The negative strain values can be explained by individual grains which are oriented in such a manner that yield easily occurs and therefore they are bent or pushed by the surrounding grains due to rotations of the slip systems in the tensile direction [8, 31]. Fig. 19-6 b corresponds to rn = 607 MPa and is the last strain image directly before the fracture of the specimen. As is known from ductile macrospecimens, a very pronounced necking with local stress and strain concentration occurs. Light reflections in the necking zone lead to a loss of local gray-scale patterns and therefore to a loss of strain facets which would result in a discontinuous stress–strain curve. Under the assumption of a continuous stress–strain curve, the loss of the strain facets was leveled out. The strains along the section line in Fig. 19-6 are presented in Fig. 19-7 for several nominal stresses. Lines a–c represent the strain development in the elastic re-

Fig. 19-7 Section lines through strain fields corresponding to Fig. 19-6 at different stress levels: a–d elastic regime; d–h, elasto-plastic regime

19.5 Results

gime and lines d–h those in the elasto-plastic regime. By tracking the zones with different compliances, typical multi-grain behavior [8, 9] can be detected. For example, at Y = 0.18 mm a relatively stiff area can be seen: in the purely elastic regime only the whole strain level increases with increasing elongation. In the elasto-plastic regime, this zone starts to plastify and exceeds even the strain values of the neighboring grains. This observation can be explained by the well-known grain boundary hardening effect, i.e. dislocations in a deformable grain pile up at a grain boundary next to a grain with a stiffer orientation. This local stress concentration is also induced into the neighboring grain and when it is high enough dislocations start to moving there also [8]. In the elasto-plastic regime, several strain maxima along the section line occur. However, already in the section line e at rn = 602 MPa the two maxima at Y = 0.2 and 0.35 mm show the area of failure. The absolute maximum of the last strain field relating to section line h amounts to 50.1%, which is distinctly higher than the maximum of section line h. The fracture surface of this specimen is shown in Fig. 19-8 a. Very conspicuous are the large deformations and the very uneven fracture surface. The size of the gap in the fracture surface in which a pore with a diameter of 10 lm occurs corresponds well with the mean grain size of 64 lm. This gap and the uneven fracture surface itself with large facets indicate additionally the strong influence of the grain anisotropy on the deformation behavior. The stress–strain curves and the mechanical parameters of batch F in Fig. 19-9 and Table 19-6 show a slightly lower scatter of the material behavior and higher strength values compared with batch E. The difference in the mean yield strengths ry of batches E and F amounts to nearly 200 MPa, and the difference in the ultimate tensile strengths is also *200 MPa. This could be expected owing to the lower temperature of the mold and the resulting smaller grain size of 35.1 lm. The scatter of the Young’s moduli E and of the elongations to fracture A is relatively low compared with batch E. The higher A values for batch F than batch E could be explained by the smaller grain size, which results in a larger number of activated slip systems. The fracture surface of the specimen with rm = 818 MPa is shown in Fig. 19-8 b, which shows dimples on the fracture surface itself and necking that is obviously less extensive than that of the batch E specimen. Owing to a slightly higher num-

Fig. 19-8 Fracture surfaces of (a) the batch E specimen with rm = 658 MPa and (b) the batch F specimen with rm = 818 MPa

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19 Analysis of Microstructure, Surface Topography and Mechanical Properties

Fig. 19-9 Stress–strain curves for batch F (VPC, Tm = 700 8C)

Table 19-6 Mechanical properties of batch F (VPC, Tm = 700 8C)

ry (MPa)

m

E (GPa)

Min.

Max.

Mean

Min.

Max.

Mean

Min.

Max.

Mean

497

607

536

101

105

103

0.24

0.46

0.35

rf (MPa)

rm (MPa)

A (%)

Min.

Max.

Mean

Min.

Max.

Mean

Min.

Max.

Mean

749

858

806

647

795

730

9.61

13.43

11.00

ber of grains in the cross section, the necking of the batch F specimen is not dominated by individual grains. However, the number of grains is still relatively low so that the deformation behavior itself is still heavily influenced by grain anisotropy, which could be seen at the scatter of the ry and rm values.

19.5.2.2 Microtensile Specimens Produced by CC The microtensile specimens of batch G show a very homogeneous microstructure (Fig. 19-10, left) and an average grain size of 10.7 lm. It was considered that the reason for the smaller grain size compared with that of the microsticks produced at the same temperature of the ceramic mold (batch C) is the supercooling of the metal melt in the cold runner. Therefore, in the next step the runner was preheated to Tr = 700 8C just like the ceramic mold (batch H). The microstructure of these microtensile specimens (Fig. 19-10, right) is very similar to that of batch G. The grain size reached (11.2 lm) is slightly larger than that in batch G. This means that the temperature of the runner plays only a subordi-

19.5 Results

Fig. 19-10 Microstructure of the microtensile specimens produced by CC at Tm = 700 8C (batch G) and at Tm = 700 8C/ Tr = 700 8C (batch H)

nate role in the grain size. The porosity of the microtensile specimens of batch G was determined as 1.5%, which is falsified by the presence of the two large pores in the specimen. Excluding these, the porosity tends to 0%, which therefore was included in Table 19-4. The porosity of batch H shows a very low value of 0.1%. Because of the higher pressure (20–25 bar) used for filling by CC, the metal melt penetrates more strongly in the ceramic mold compared with the VPC specimens of batch E, which causes the high peak-to-valley roughness of 8.7 lm in batch G (Fig. 19-11, left) and 7.5 lm in batch H instead of 4.3 lm after VPC (Fig. 19-11, right). The stress–strain curves of five specimens are shown for batch G in Fig. 19-12 and the related mechanical properties are given in Table 19-7. The stress–strain curves and the mechanical properties show a relatively large scatter of the material behavior. The mean yield strength ry, the Young’s moduli E and the ultimate tensile strengths rm are considerably higher than those of batch E specimens and correspond very well to those of batch F specimens which were produced by VPC. De-

Fig. 19-11 Surface topography of the microtensile specimens produced by CC at Tm = 700 8C (batch G, left) and at Tm = 700 8C/Tr = 700 8C (batch H, right)

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19 Analysis of Microstructure, Surface Topography and Mechanical Properties

Fig. 19-12 Stress–strain curves for batch G (CC, Tm = 700 8C, Tr = room temperature)

Table 19-7 Mechanical properties of batch G (CC, Tm = 700 8C, Tr = room temperature)

ry (MPa)

m

E (GPa)

Min.

Max.

Mean

Min.

Max.

Mean

Min.

Max.

Mean

450

609

518

91

121

106

0.14

0.35

0.30

rf (MPa)

rm (MPa)

A (%)

Min.

Max.

Mean

Min.

Max.

Mean

Min.

Max.

Mean

728

952

824

647

908

785

3.77

9.95

6.01

Fig. 19-13 Typical fracture surface of the batch G specimens

spite an considerable difference in the grain sizes of batches F and G, the ry and rm values are comparable. Therefore, it can be concluded that the superlattice has been modified owing to the production process itself. The mean value of the elongation to fracture A is lower. In Fig. 19-13, a typical fracture surface of the batch G specimens is shown. In contrast to the fracture surfaces of batch E and F specimens, this

19.5 Results

fracture surface shows obviously higher porosity, which could explain the relative large scatter of the ry, E, rm and A values. Furthermore, the fracture surface shows much more dimples and less deformations due to necking. Additionally, the scatter of the mechanical properties could still be influenced by anisotropy. The stress–strain curves and the mechanical parameters of batch H in Fig. 19-14 and Table 19-8 show a comparable material behavior to batch G. The tensile curve of the individual specimen which had the mechanical properties ry = 245 MPa, E = 92 GPa, m = 0.27, rm = 445 MPa and A = 1.08% failed owing to the large notch effect of a large pore with dimensions of about 30 ´ 50 lm and is therefore not considered in Table 19-8. The yield strengths ry of the other specimens of batch H have a mean value of 443 MPa, which is 75 MPa lower than the mean value of batch G. In contrast, the mean ultimate tensile strength rm = 818 MPa and the mean elongation to fracture A = 6.67% are comparable to the mean values for batch G. The fracture surfaces correspond to that of batch

Fig. 19-14 Stress–strain curves for batch H (CC, Tm = 700 8C, Tr = 700 8C) Table 19-8 Mechanical properties of batch H (CC, Tm = 700 8C, Tr = 700 8C)

ry (MPa)

m

E (GPa)

Min.

Max.

Mean

Min.

Max.

Mean

Min.

Max.

Mean

325

511

443

83

140

118

0.35

0.39

0.37

rf (MPa)

rm (MPa)

A (%)

Min.

Max.

Mean

Min.

Max.

Mean

Min.

Max.

Mean

702

882

818

672

845

785

5.45

7.79

6.67

541

542

19 Analysis of Microstructure, Surface Topography and Mechanical Properties

Fig. 19-15 Microstructure (left) and surface topography (right) of the microtensile specimens produced by CC at Tm = 1000 8C (batch I)

G shown in Fig. 19-13 and are therefore not presented here. Overall the mechanical properties correspond well with those of batch G, which means that the influence of the preheated runner can be neglected, as already concluded from the analysis of the microstructure. The microtensile specimens of batch I show a homogeneous microstructure with a grain size of 10.2 lm (Fig. 19-15, left). This is an enormous difference in comparison with the grain size of the microtensile specimens (64.0 lm) which were produced by VPC using the ceramic mold at the same temperature (batch E). As already mentioned in this section, the temperature of the runner is of only secondary interest because of the slight difference in the grain size when using a preheated and a cold runner. A cause of this reduction of the grain size by the CC process may be the temperature of the ceramic mold. The preparation process for the CC, which includes evacuation of the equipment and flushing with argon, takes 3 min. It is assumed that during this time the ceramic mold is cooled and the metal melt is supercooled faster than expected. This will be investigated in further work. The porosity of the microtensile specimens of the batch I was determined as 0.7%. The microspecimens show a relatively rough surface topography (Fig. 1915, right) and a peak-to-valley roughness of 7.9 lm because of the higher pressure by the CC process. The stress–strain curves and the mechanical parameters of batch I in Fig. 1916 and Table 19-9 show a similar material behavior to batches G and H. In comparison with batch E, the yield strengths and the ultimate tensile strengths are again considerably higher. As already described in Sections 19.5.2.1 and this section, this is probably due to the modified superlattice and the smaller grain size. The Young’s moduli E show a lower mean value and a larger scatter compared with batches G and H. The Poisson’s ratios m of batch I are equivalent to batch G and comparable to all other batches. The ultimate tensile strengths rm resulted in a mean value of 776 MPa and a relatively large scatter interval of 240 MPa. The mean elongation to fracture A is comparable to those of batches G and H. However, the scatter of the A values of batch I is distinctly higher. The specimen which has the lowest A value of 3.54% failed owing to large

19.5 Results

Fig. 19-16 Stress–strain curves for batch I (CC, Tm = 1000 8C, Tr = room temperature) Table 19-9 Mechanical properties of batch I (CC, Tm = 1000 8C, Tr = room temperature)

ry (MPa)

m

E (GPa)

Min.

Max.

Mean

Min.

Max.

Mean

Min.

Max.

Mean

488

543

521

63

119

94

0.23

0.41

0.33

rf (MPa)

rm (MPa)

A (%)

Min.

Max.

Mean

Min.

Max.

Mean

Min.

Max.

Mean

658

898

776

636

813

730

3.54

11.6

7.25

pores with a cross sections of 20 ´ 30 lm. The fracture surfaces of the other specimens also showed pores which are smaller but lead overall to large scatter of the mechanical properties. Additionally, an influence of anisotropy cannot be excluded. Owing to the comparability of the microstructure and the mechanical properties of batches G, H and I, it can be concluded that the supercooling of the ceramic mold of all three CC production processes must have been comparable and considerably higher than that of the VPC batches E and F.

19.5.2.3 Comparison of Microtensile Specimens Produced by CC and by VPC In order to compare the strengths of the investigated microspecimens according to their microstructures and mechanical properties, a formal Hall–Petch diagram is presented in Fig. 19-17. The VPC microspecimens reveal, in addition to the already described hardening by ordering due to fine coherent CuAu I precipitates, an influence of the grain size: Smaller grains result in higher strength

543

544

19 Analysis of Microstructure, Surface Topography and Mechanical Properties

Fig. 19-17 Hall–Petch diagram of the investigated tensile specimens of batches E and F (VPC), G–I (CC) and J (CC, Tm = 1000 8C, subsequently annealed for 5 h at 800 8C)

values, which are in accordance with the aforementioned Hall–Petch-effect. The CC microspecimens as cast each have similar strength values and grain sizes despite the different cooling conditions. This can be deduced from the observation that there is no major influence of the mold and the runner temperature on the microstructure and the mechanical properties, which results in similar material states. In comparison with the VPC specimens, the smaller grain sizes of the CC specimens can be deduced from faster solidification of the melt. Owing to the smaller grain sizes of the CC specimens, a higher strength level would be expected. The lower strength level can probably be ascribed to the minor formation in the CC specimens of the superstructure that is clearly apparent in the VPC specimens. The formation of the superstructure is dependent on the cooling rate in the range 300–400 8C. In order to study the development of the superstructure in direct correlation with the cooling conditions, further studies have to be performed.

19.5.2.4 Microtensile Specimens Produced by CC with Additional Annealing at 800 8C for 5 h In comparison with the as-cast microtensile specimens of batch I, identically cast microtensile specimens which were additionally annealed at 800 8C for 5 h (batch J) show inhomogeneous grain growth (Fig. 19-18, left). The average grain size is increased from 10.2 to 35.9 lm. The porosity of the annealed microtensile specimens tends to 0%. The surface topography of batch J shown in Fig. 19-18, right, looks relatively smooth compared with the other microtensile specimens produced by CC but shows a peak-to-valley roughness of 6.4 lm. This value could be explained by the relatively large scatter of the roughness values along the whole gauge length.

19.5 Results

Fig. 19-18 Microstructure (left) and surface topography (right) of the microtensile specimen produced by CC at Tm = 1000 8C and additional annealing at 800 8C for 5 h (J)

Fig. 19-19 Diffraction patterns of the microtensile specimens produced by CC at Tm = 1000 8C (I) and at Tm = 1000 8C with additional annealing at 800 8C for 5 h (J)

For a detailed analysis of the microstructure, the annealed microtensile specimens were also investigated using TEM. In Fig. 19-19, the diffraction patterns of the batches I and J are presented. In both of them the typical formation of a CuAu I superstructure is shown. The diffraction pattern of the annealed microtensile specimens shows a clearly more distinct superstructure formation. In order to achieve complete dissolution of the superstructure, the microtensile specimens were annealed at 800 8C. Because of the subsequent slow cooling from annealing temperature to room temperature, the formation of the CuAu I superstructure was promoted. This is also confirmed by the dark field images (Fig. 19-20) which were formed using a longish secondary weak reflex of the superstructure. In both of the dark field images the formation of coherent CuAu I ordered platelets is shown. These platelets demonstrate the interface boundary between ordered domains. It can clearly be seen that the CuAu I platelets of the annealed microtensile spec-

545

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19 Analysis of Microstructure, Surface Topography and Mechanical Properties

imens are longer and more massive (Fig. 19-20, right) than in the microtensile specimens which were not annealed (Fig. 19-20, left). The heat treatment resulting in the modified morphology of the superstructure of batch J leads to the stress–strain curves and the mechanical properties of the two batch J specimens shown in Fig. 19-21 and Table 19-10. Compared with batch I, distinctly higher yield strengths, ultimate tensile strengths and Young’s moduli and lower elongations to fracture are found. Owing to the grain size and the low or vanishing porosity, the substantial differences in the ry and E values of the two batch J specimens indicates an influence of the grain anisotropy. As could be expected, the fracture surfaces of the specimens show rela-

Fig. 19-20 Dark field images of the microtensile specimens produced by CC at Tm = 1000 8C (I) and at Tm = 1000 8C with additional annealing at 800 8C for 5 h (J)

Fig. 19-21 Stress–strain curves for batch J (CC, Tm = 1000 8C, Tr = room temperature, annealed for 5 h at 800 8C)

19.5 Results Table 19-10 Mechanical properties of batch J (CC, Tm = 1000 8C, Tr = room temperature, annealed for 5 h at 800 8C)

ry (MPa)

m

E (GPa)

Min.

Max.

Mean

Min.

Max.

Mean

Min.

Max.

Mean

629

802

716

97

135

116

0.31

0.38

0.35

rf (MPa)

rm (MPa)

A (%)

Min.

Max.

Mean

Min.

Max.

Mean

Min.

Max.

Mean

938

1072

1005

933

1039

986

4.35

4.51

4.43

tively low deformations, are relatively smooth and show dimples all over the cross section. The mean strength values of the annealed CC microspecimens which are also presented in the Hall–Petch diagram in Fig. 19-17 are distinctly higher compared with the other CC specimens, despite a larger grain size. This is caused by the strong formation of the superstructure, which leads even to higher strength values than in VPC specimens with similar grain sizes. 19.5.3 Microstructure and Mechanical Properties of Microbending Specimens

The microbending specimens of batch K, which were produced by VPC at Tm = 700 8C, subsequently water quenched and then annealed for 20–30 min at 400 8C, show a homogeneous microstructure (Fig. 19-22, left) with grain sizes between 20 and 50 lm. As shown in Table 19-11, the mean grain size is 41.4 lm. The grain boundaries are decorated with a secondary phase. In additional studies using TEM, it was found that this phase contains subgrains, twins and dislocations. The porosity of the microbending specimens of batch K was determined as 0.5%. The microspecimens show a relatively smooth surface topography (Fig. 19-22, right) and have a peak-to-valley roughness of 3.6 lm.

19.5.3.1 Quasi-static Micro-three-point Bending The yield strengths of five bending specimens of batch K vary in the range 824– 1075 MPa, with a mean value of 926 MPa. Owing to the ductility of the investigated gold-based dental alloy, the bending specimens could not be deflected until fracture occurred. The quasi-static bending tests were also accompanied by optical strain measurements as explained in the experimental details. Fig. 19-23 shows a photograph of a loaded specimen, obtained by the strain measurement

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19 Analysis of Microstructure, Surface Topography and Mechanical Properties

Fig. 19-22 Microstructure and surface topography of the batch K microbending specimens produced by VPC at Tm = 700 8C, subsequently water quenched and then annealed for 20– 30 min at 400 8C

Table 19-11 Casting technique, cooling conditions, material type of the mold, grain size and porosity of batch K VPC microbending specimens

Batch

Casting technique

Cooling conditions

Material type of mold

K

VPC

Tm=700 8C, water Deguvest quenched, annealed at 400 8C, 20–30 min

Grain size (lm)

Porosity (%)

Peak-tovalley roughness (lm)

41.4

0.5

3.6

Fig. 19-23 Loaded microbending specimen and corresponding strain field on the surface

19.5 Results

Fig. 19-24 Development of the local strain in the longitudinal direction corresponding to the section lines given in Fig. 19-23

system, and the corresponding distribution of the strain in the longitudinal direction. The strains on the sides loaded in tension or compression show distinct differences from the material behavior that would be expected for quasi-isotropic material in three-point bending. The dimensions of the areas, showing local strain concentrations, correlate relatively well with the average grain size of 41.4 lm. Owing to this and the relatively small number of grains in the cross section, grain anisotropy can be considered as a reason for the observed strain distribution and for the relatively large scatter of the yield strengths. The dashed lines in the strain image in Fig. 19-23 indicate the positions of the sections which are given for different deflections d in Fig. 19-24. The longitudinal strain of the Y-section in Fig. 19-24, left, shows relatively linear strain distributions. The section in the longitudinal direction along the specimen side loaded in tension in Fig. 19-24, right, shows the development of the above-mentioned strain concentrations with increasing deflection d. The transition from elastic to elasto-plastic behavior occurs between the deflection lines a and b at a fictitious bending stress rb of about 900 MPa. However, the evolution of local strain concentrations already starts in the elastic regime and increases with increasing deflection at two or three points more explicitly than in the neighboring surface regions. The region which is in the area of the maximum bending moment shows, in contrast to the regions to the right and left, distinctly lower strains, which could be explained by a stiffer orientation of the grains located there.

19.5.3.2 Cyclic Micro-three-point Bending The cyclic micro-three-point bending tests were performed with force control with an elastically calculated mean stress rm of 700 MPa and superimposed fictitious stress amplitudes ra = 300 between 200 and 650 MPa. The lifetimes obtained from all fatigue tests result in the Wöhler diagram presented in Fig. 19-25 [27].

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19 Analysis of Microstructure, Surface Topography and Mechanical Properties

Fig. 19-25 Wöhler diagram at rm = 700 MPa for the gold-based alloy Stabilor G (batch K)

The respective cyclic deformation curves of the displacement amplitudes sa and of the mean displacements sm both versus the number of cycles N at the fictitious stress amplitudes ra = 300, 400, 500 and 600 MPa are shown in Fig. 19-26 a. During the first cycle, large plastic deformations occur, which increase with increasing ra. Subsequently, continuous but weak cyclic hardening is observed. The rate of this cyclic hardening is the larger the higher is the load amplitude. After the first cycle, the mean displacement remains nearly constant for ra  500 MPa. At ra = 600 MPa, continuous but very weak cyclic creep is observed. Despite the relatively large scatter of the yield strengths, as mentioned above, the scatter of the lifetimes is below a factor of 4. This observation can be explained by comparing the sa–N courses of two identically loaded specimens,

Fig. 19-26 Cyclic deformation behavior of Stabilor G at  = 700 MPa rm

19.5 Results

e.g. rm = 700 MPa and ra = 500 MPam as shown in Fig. 19-26 b. In the first cycle, specimen 1 (solid curve) shows a clearly higher plastic strain amplitude and a corresponding increase in the mean displacement. Up to N & 1000, this specimen shows more pronounced cyclic hardening than the other sample. However, at N ³ 1000, the displacement amplitudes of both specimens remain at nearly identical values until failure occurs. Consequently, during more than 90% of the lifetime, identical plastic strain amplitudes occur in both specimens, which leads to a much less pronounced difference in the number of cycles to failure NF than would be expected from the cyclic deformation behavior during the first cycles. Fig. 19-27 shows SEM images of a fracture surface after cyclic loading with rm = 700 MPa and ra = 400 MPa. The fatigue crack initiates and propagates along slip planes with an angle of about 458 with respect to the normal stress in the region of the surface that is loaded in tension (stage I of crack propagation, see top of Fig. 19-27 a). Subsequently, the crack direction becomes nearly perpendicular with respect to the direction of the longitudinal tensile stress (stage II of crack propagation). In the respective region of the fracture surface, slip bands become visible with a distance that increases with increasing crack length (see Fig. 19-27 b). The remaining fracture surface shows dimples which are typical for the rupture of ductile materials.

Fig. 19-27 Fracture surface of a cyclically loaded Stabilor G microbending specimen with slip planes in the stage I zone of crack propagation [top of (a)] and slip bands in the stage II zone of crack propagation (b)

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19 Analysis of Microstructure, Surface Topography and Mechanical Properties

19.6 Conclusion and Outlook

In this chapter, investigations on correlations between process parameters, state of the microspecimens and their mechanical properties are presented exemplarily for microcasting of the gold-based alloy Stabilor G. Microcasting was performed by vacuum pressure casting (VPC) and centrifugal casting (CC) with different cooling conditions adjusted by modifying the temperatures of the mold and the runner. The microstructures after VPC show a reduction of dendritic structures and increasing grain sizes with increasing mold temperature. The porosity and roughness are low. After CC, no significant influence of the temperature of the preheated mold and that of the runner on the grain size could be found, which is probably due to comparable supercooling of the mold or runner, respectively, during the CC process. The porosity is slightly higher and the roughness increases considerably because of the higher casting pressure. The mechanical properties of quasi-static tensile tests after VPC show severe scatter because of the relatively large grain sizes, which leads to effects of elastic and plastic anisotropy. This could be shown by locally resolved strain fields and fits very well with previous investigations of the deformation behavior of Stabilor G [27]. After CC, despite the smaller grain sizes, the scatter of the mechanical properties of quasi-static tests is still very high. Although anisotropy effects or the influence of textures cannot be completely excluded, this may be deduced from a generally higher porosity or single large pores. Whereas the grain sizes of the CC specimens are much lower than those of the VPC specimens, their strengths are similar to those of the VPC specimens with the smaller grain size. This is due to the suppression of the superstructure, which lowers the strength. An additional annealing treatment of CC specimens allows the formation of the superstructure but leads to larger grain sizes. Overall, this results in a distinct increase in the strength. Fatigue tests on VPC bending specimens show a weak cyclic hardening during cyclic loading. In spite of a relatively large scatter of the yield strength, the scatter of the fatigue lifetimes is relatively low.

19.7 References G. Baumeister, K. Mueller, R. Ruprecht, J. Hausselt, J. Microsyst. Technol. 2002, 8, 105–108. 2 K. Yasuda, Gold Bull. 1987, 20 (4), 90– 103. 3 D. Ott, Handbook on Casting and Other Defects in Gold Jewellery Manufacture; London: World Gold Council, 1997, p. 77. 1

E. O. Hall, Proc. Phys. Soc. London B 1951, 64, 747. 5 N. J. Petch, The cleavage strength of polycrystals, J. Iron Steel Inst. 1953, 174, 25. 6 OMG, Edelmetalltaschenbuch; Hanau: Giesel Verlag, 2001. 4

19.7 References 7 H. Kuchling, Taschenbuch der Physik,

8

9 10

11 12

13

14

15

16 17

18

13th edn.; Leipzig: Verlag Harri Deutsch, 1991. R. W. K. Honeycombe, The Plastic Deformation of Metals, 2nd edn.; London: Edward Arnold, 1984. M. Sachtleber, Z. Zhao, D. Raabe, Mater. Sci. Eng. A 2002, 336, 81–87. S. Zaefferer, J.-C. Kuo, Z. Zhao, M. Winning, D. Raabe, Acta Mater. 2003, 51, 4719–4735. G. Baumeister, R. Ruprecht, J. Hausselt, J. Microsyst. Technol. 2004, 10, 261–264. B. Kasanická, M. Fotouhi, V. Schulze, D. Gerthsen, D. Löhe, ‘Transmissions-elektronenmikroskopische Untersuchungen der AuAgCu-Dentallegierung Stabilor G, Praktische Metallographie,’ in preparation. B. Kagerer, R. Brodmann, J. Valentin, J. Filzek, U. Popp, Proc. SPIE 2002, 4773, 52–62. ‘Geometrische Produktspezifikationen, Oberflächenbeschaffenheit, Tastschnittverfahren, Regeln und Verfahren für die Beurteilung der Oberflächenbeschaffenheit,’ DIN EN ISO 4288; Berlin: Beuth Verlag, 1998. E. Mazza, Experimental Methods for the Characterization of Micrometer-sized Mechanical Structures; Dissertation, ETH Zürich, 1997. A. Ilzhöfer, H. Schneider, Ch. Tsakmakis, J. Microsyst. Technol. 1997, 4, 46–50. Y. Higo, K. Takashima, M. Shimojo, S. Sugiura, M. V. Swain, in: Conf. Proc. Micro Mat 2000, Berlin; 2000, pp. 997–1000. U. Rabe, S. Amelio, M. Kopycinska, S. Hirsekorn, M. Kempf, M. Göken, W. Ar-

19 20 21

22

23

24 25

26 27

28

29

30

31

nold, Surf. Interface Anal. 2002, 33, 65– 70. N. Huber, W. D. Nix, H. Gao, Proc. R. Soc. London A 2002, 458, 1593–1620. N. Huber, E. Tyulyukovskiy, J. Mater. Res. 2004, 19, 101–113. K. Obergfell, T. Beck, V. Schulze, D. Löhe, MP Materialprüfung, 2000, 42, 391–395. P. Ruther, W. Bacher, K. Feit, Wissenschaftliche Berichte Forschungszentrum Karlsruhe (FZKA); 1997, No.5986. W. N. Sharpe Jr., ‘Mechanical properties of MEMS materials,’ in: The MEMS Handbook; Boca Raton, FL: CRC Press, 2001, Chapter 3, pp. 3-1–3-33. W. N. Sharpe Jr., Proc. Soc. Exp. Mech. 2003, 43, 228–237. M. Auhorn, B. Kasanická, T. Beck, V. Schulze, D. Löhe, Z. Metallkd. 2003, 94, 599–605. M. Auhorn, T. Beck, V. Schulze, D. Löhe, J. Microsyst.Technol. 2004, 10, 489–492. M. Auhorn, T. Beck, V. Schulze, D. Löhe, in: Conf. Proc. Fatigue 2002, Stockholm, A. F. Blom (ed.); 2002, Vol. 4/5, pp. 2641–2648. GOM, Aramis-Handbuch, Verformungsmessung nach dem Rasterverfahren; Aramis 4.7.4–2, Braunschweig, 2003. B. Kasanická, V. Schulze, D. Gerthsen, D. Löhe, in: Sonderbände der Praktischen Metallographie 35, G. Petzow (ed.); 2003, pp. 451–456. H. Berns, Metallurgischer Größeneinfluss; VDI-Berichte No. 354, Düsseldorf: VDI, 1979, pp. 27–38. D. Raabe, Z. Zhao, S.-J. Park, F. Roters, Acta Mater. 2002, 50, 421–440.

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20 Microstructure, Surface Topography and Mechanical Properties of Molded ZrO2 Microspecimens M. Auhorn, B. Kasanická, T. Beck, V. Schulze, D. Löhe, Institute of Material Science and Engineering I, University of Karlsruhe (TH), Germany Abstract Several batches of zirconia (ZrO2) three-point bending specimens with the dimensions 1.2 ´ 0.2 ´ 0.2 mm produced by low- or high-pressure injection molding were investigated with respect to their microstructure, surface topography, residual stress and behavior under quasi-static loading. According to the Weibull size effect, the bending strengths of the microspecimens are higher than those of comparable macrospecimens. The microspecimens examined have characteristic bending strengths up to 3181 MPa, the strength being heavily influenced by surface quality, edge geometry and porosity. Additionally, these failure influencing factors were partly simulated by finite element method calculations.

Keywords ZrO2; microstructure; surface topography; edge radius; micro-three-point bending

20.1 20.2 20.2.1

20.2.2 20.3 20.4 20.4.1

20.4.2 20.4.3 20.5 20.6

Introduction 556 Test Material and Production of Specimens

556

ZrO2 Specimens Produced by Low-pressure Injection Molding (LPIM) 556 ZrO2 Specimens Produced by High-pressure Injection Molding (HPIM) 558 Experimental Details 558 Results and Discussion 562

Characterization of Microstructures, Surface Topographies and Residual Stresses 562 Determination of Mechanical Properties and Influencing Factors 565 Simulation of Influencing Factors on the Bending Strength 573 Conclusions and Outlook References 578

576

Advanced Micro and Nanosystems Vol. 4. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31493-8

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20 Microstructure, Surface Topography and Mechanical Properties of Molded ZrO2 Microspecimens

20.1 Introduction

One problem in designing parts in microtechnology is the lack of reliable materials data. The existing material properties of macroscopic specimens normally cannot be used owing to the neglect of some important aspects of the ‘microworld,’ e.g. the surface-to-volume ratio is much higher, which results in a greater influence of surface defects on the material strength. Additionally, the number of grains in relevant cross sections is smaller and for this reason effects of anisotropy become of major interest. Therefore, one of the aims of the SFB 499 is to formulate a microspecific dimensioning concept for designing microparts [1–3]. In order to identify relevant process parameters, it is also necessary to investigate their influence on the properties of microspecimens from the respective microproduction processes, here low-pressure [4–6] and high-pressure injection molding [7, 8] with ZrO2 powder. The present study was aimed at investigating the microstructure, the residual stresses, the surface topography and the mechanical properties of ZrO2 microbending specimens. The influencing factors on the bending strengths were determined and partly simulated by finite element method (FEM) calculations. The results allow important aspects of a dimensioning concept for microparts made of engineering ceramics to be formulated.

20.2 Test Material and Production of Specimens

The material examined was sintered polycrystalline ZrO2 partially stabilized with 3 mol% Y2O3 and with an additional small content of Al2O3 as a sintering aid. The microbending specimens investigated were produced by micro-ceramic injection molding at the Institut für Materialforschung III of the Forschungszentrum Karlsruhe. Within the framework of the SFB 499, two alternative processing routes are applied for the manufacture of ceramic microparts: low-pressure injection molding (LPIM) and high-pressure injection molding (HPIM). The principles, specific advantages and limitations of these shaping methods are discussed in detail in Chapter 12. 20.2.1 ZrO2 Specimens Produced by Low-pressure Injection Molding (LPIM)

An array of 225 rectangular pillars was generated by micromilling of a brass block. This structure served as master model for the replication by LPIM in ZrO2. For the replication, a mold was prepared which was made of highly flexible silicone rubber. As very low viscosities of the powder/binder compound (‘feedstock’) are a characteristic feature of the LPIM process, only small forces are necessary for injection, allowing distortion-free filling of the soft silicone rubber mold. The mold filling in LPIM is typically performed by a piston or by

20.2 Test Material and Production of Specimens

air pressure. However, for the investigations presented here only a small number of arrays, each containing a large number of specimens, were required. These arrays were prepared from low-viscosity LPIM feedstocks by a simple manual casting-like process. In contrast to the machined process, where the feedstock is injected into an evacuated cavity, the removal of entrapped air is performed after the filling process in a heated vacuum chamber. Therefore, the residence time of the molten feedstock in the silicone rubber mold is prolonged for the manual filling process. Feedstocks were prepared with a solid content of 50–52 vol.% ZrO2 powder from TOSOH (TZ-3YS-E, 0.25 wt.% Al2O3) and a paraffin binder. Suitable dispersants were added to the mixture to decrease the feedstock viscosity to the desired level. For the mold filling, the feedstock was heated to 80–100 8C. After molding, the ceramic mass was degassed in a vacuum chamber to remove air inclusions. After cooling of the mold, the solidified array was removed manually from the silicone rubber mold. Then the remaining so-called green body was subjected to thermal debinding in an air atmosphere. As listed in Table 20-1, six batches were produced by LPIM with varying feedstock compositions, viscosities and debinding conditions. The debinding of batches A and B was performed at 500 8C with a heating rate of 0.15 K/min and several dwell times at different temperatures. The final dwell time of the debinding was 300 min. Subsequently the specimens were cooled to room temperature in the debinding furnace. The debinding of batch C was realized at 400 8C with a heating rate of 0.25 K/min and a dwell time of 10 min. In comparison with batch C, batch D had an additional dwell time of 5 h at 180 8C. For batches E and F, debinding was performed at 400 8C with a dwell time of 10 min. Heating from room tem-

Table 20-1 Production parameters of the LPIM (batches A–F) and HPIM (batch G) specimens

Parameter

Molding Feedstock: vol.-% ZrO2 Shear viscosity at 85 8C a), 190 8C b) (shear rate 100 s–1) (Pa s) Debinding: LPIM thermally in air, HPIM catalytic in N2 + HNO3 Sintering

Batch A

B

C

D

LPIM I: 50

II: 51.5

III: 52

IV: 52

HPIM V: 50

4.5

7.5

26

21

500

0.25 K/min, 400 8C, 10 min

4 K/min, 110 8C, 20 min

0.15 K/min, 500 8C, 300 min

0.25 K/min, 400 8C, 10 min; D: additional 5 h at 180 8C 3 K/min, TS = 1500 8C, tS = 1 h

Working temperature of feedstock for

a)

LPIM and

b)

HPIM.

E

F

G

557

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20 Microstructure, Surface Topography and Mechanical Properties of Molded ZrO2 Microspecimens

perature was performed at *0.25 K/min with several dwell times at different temperatures. Finally, a sintering treatment was performed in a box furnace in an ambient air atmosphere at 1500 8C for 60 min. By this procedure, arrays of bending specimens standing on a base plate were obtained [4–6]. 20.2.2 ZrO2 Specimens Produced by High-pressure Injection Molding (HPIM)

For the production of the high-pressure injection molded microspecimens (Table 20-1, batch G), the commercially available feedstock Catamold TZP-A from BASF with 50 vol.% solid content of ZrO2 powder was used. The first step in the micro-high-pressure injection molding process [9, 10] was the production of a green compact by injecting the heated feedstock in a heated and evacuated microstructured mold which was produced by micromilling. The maximum pressure in the injection molding machine which occurred during the process was 180 bar. After injecting the feedstock, the mold was cooled and the part was taken out. After this process, the green compact was subjected to catalytic debinding in an N2/HNO3 atmosphere: the specimens were heated to 110 8C at 4 K/min and kept there for 20 min. In order to obtain dense microparts, these brown bodies were sintered for 60 min at 1500 8C in an air atmosphere [7, 8].

20.3 Experimental Details

For the analysis of the microstructure, the specimens produced by low- or highpressure injection molding were thermally etched for 1 h at 1350 8C. The microstructure was examined using light microscopy (Leica Aristomet) and scanning electron microscopy (LEO Gemini). The grain size was determined by applying the intercepted-segment method. The porosity was measured by using the binary method, i.e. by optical separation of the gray-scale spectrum using computeraided algorithms. The potential sources of errors of the binary method are described in [11]. Furthermore, the surface topographies of the specimens and the related peakto-valley heights Rt were measured at the surfaces that were subsequently loaded in tension by using a confocal white-light microscope (NanoFocus) [12]. The Rt values could only be determined in longitudinal specimen direction according to DIN EN ISO 4288 [13], which recommends a minimum measuring length of 0.48 mm. Additionally, this standard suggests a cut-off wavelength of 250 lm in order to distinguish the roughness from the waviness. Generally, it cannot be completely excluded that some of the widespread surface defects of the specimens are recognized as parts of the waviness profile, which leads to slightly underestimated Rt values in some cases. The analysis of residual stresses was carried out using parallel synchrotron radiation at the HASYLAB beamline G3/DESY, Hamburg, Germany. The experi-

20.3 Experimental Details

mental set-up at G3 consists of a four-circle MAXIM diffractometer and a microchannel plate (MCP) as a collimator in front of a x-ray sensitive CCD camera (Fig. 20-1). This arrangement offers the possibility of obtaining diffraction data for the entire microcomponent with high spatial resolution [14, 15], which is necessary in view of the very small dimensions of the microspecimens. Here, the residual stress analysis was performed in the X-mode. A wavelength of k = 0.1544 nm was used to determine the lattice strain of the {313}-planes. Bragg peaks were measured in the range 81.2408 £ 2h £ 82.5528 by D2h = 0.0168 using 17 different angles of incidence 8.448 £ X £ 72.448 with a step size in x of 48. The different angles of incidence used led to an effective w range of –328 £ w £ 328. Bragg peaks were registered along the longer side of the microbending specimens. The results were evaluated by using the sin2w method [16]. A spatial resolution of 150 ´ 150 lm2 on the specimen was reached. This resolution arises from the parameters of the measurement used and from the necessity to have at least three pixels for the analysis of the data. For the evaluation a rectangular area in the center region of the specimen was selected. The mechanical properties of the microspecimens were determined with a special universal microtesting machine [17, 18] that was developed at the Institut für Werkstoffkunde I of the Universität Karlsruhe (TH). The machine has an electrodynamic actuator and allows tests with quasi-static or cyclic loadings. It can be equipped with microtensile clamps or with a micro-three-point bending device. A detailed description of the microtensile test and a short overview of micromechanical testing in general is given in Chapter 19. The bending supports were produced by micro-electrodischarge machining (EDM). The lower supports have a distance of 0.8 mm and a radius of 50 lm. The upper support has a radius of 100 lm. The load capacity is 50 N. The force is measured with a commercially available 100 N load-cell. The displacement of the lower support is determined by a capacitive displacement-measuring system. All quasi-static experiments were performed with displacement control by increasing the displacement linearly at a rate of 2 lm/s until fracture occurred [18–20].

Fig. 20-1 Scheme of the experimental setup at HASYLAB beamline G3 with a microchannel plate as collimator in front of a CCD camera

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20 Microstructure, Surface Topography and Mechanical Properties of Molded ZrO2 Microspecimens

The bending specimens, which had a length of about 1.2 mm and a cross section of approximately 0.2 ´ 0.2 mm, were tested as sintered, i.e. without any surface treatment. Before starting the mechanical tests, the dimensions of the cross sections of the individual specimens were measured using the already mentioned confocal white-light microscope. Additionally, the edge radii of selected specimens were determined in order to correlate them with their strength values after the bending tests [20]. The bending strength rb was calculated according to Equation (1) using the section modulus for bending Sb [Equation (2)] with the measured breadth b and height h of the specimen, the fracture bending moment Mb or the measured fracture force F and the distance of the lower supports l [21]: rb ˆ

Mb 3 F
l ˆ
Sb 2 b_
h2

…1†

Sb ˆ

bh2 6

…2†

The statistical analysis of the bending strength was carried out using the twoparameter Weibull theory [21]: F…rb † ˆ 1

  m  rb exp r0

…3†

The characteristic strength r0 relates to a probability of fracture of 63.2%. The Weibull exponent m represents the scatter of the strength and the reproducibility between related batches. The higher is m, the lower is the scatter and the better the reproducibility. However, owing to the dependence of m on the number of specimens tested, a corrected Weibull exponent mcorr has to be calculated with the maximum likelihood method [22]. The characteristic strength r0 depends not only on the material itself but also on the probability of critical defects in the volume V or at the surface S of a specimen and, hence, on the specimen size, resulting for the two different sample sizes 1 and 2 in the following relationships [21]: r0;1 ˆ r0;2

 m1 V2 V1

…4 a†

r0;1 ˆ r0;2

 m1 S2 S1

…4 b†

In the case of a non-uniform stress distribution, the volumes in Equation (4 a) and the surfaces in Equation (4 b) have to be replaced by effective volumes or effective surfaces. The general equations for the effective volume and the effective surface are [21]

20.3 Experimental Details

Z Veff ˆ

g m dV

…5 a†

g m dS

…5 b†

Z Seff ˆ

The geometric function g(x,y,z) represents the relation between the stress distribution r(x,y,z) and the critical bending stress at the outer fiber leading to failure. The resulting effective volume for three-point bending is given by [21, 23] Veff ;3-point bending ˆ

lbh

…6 a†

2…m ‡ 1†2

The effective surface of a three-point bending specimen is equal to [23] Seff ;3-poimt bending ˆ l

h ‡ b…m ‡ 1†

…6 b†

…m ‡ 1†2

The results of the statistical analyses were interpreted using scanning electron microscope (SEM) images of the fracture surfaces to determine the origins of failure. In practice, the strengths of microparts are influenced by several factors that cannot be separated or weighted easily. In order to obtain more detailed information for the dimensioning concept about single influences on the bending strength, it is therefore necessary to simulate the bending test with FEM calculations. By using the geometries and symmetries of the already described microbending test, the simulations were performed with a quarter model (Fig. 20-2)

Quarter model

Fig. 20-2 FEM simulation of microbending with quarter model with variation of the edge geometry (20 lm radius and 908 edge) and size of open pores (3, 6, 15, 30 lm) in the middle or at the edge of the specimen at the surface loaded in tension

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20 Microstructure, Surface Topography and Mechanical Properties of Molded ZrO2 Microspecimens

using the ABAQUS 6.4 Standard. Generally, a linear elastic and quasi-isotropic material behavior with Young’s modulus E = 210 000 MPa and Poisson’s ratio m = 0.31 was assumed. As boundary conditions the specimen was fixed at the lower support in the x, y and z directions and a displacement of Df = 15 lm at the line contact of the upper support was applied. In the present work the edge geometry, the local position of single pores at the surface loaded in tension and the size of the pores itself were varied. As shown in Fig. 20-2, the edge of the specimen was modeled as a sharp 908 edge or with a radius of 20 lm. The single pore was located at z = 0 at the surface loaded in tension, i.e. in the middle of the specimen or at the edge. The pore diameter was varied as 3, 6, 15 and 30 lm. As results of the simulations, relative stress ratios are shown along the line of maximum deflection by correlating the stress values in the longitudinal direction of a model with pores with the stress value at z = 0 of the corresponding model without pores.

20.4 Results and Discussion 20.4.1 Characterization of Microstructures, Surface Topographies and Residual Stresses

The microstructure of the bending specimens of batches A–G is exemplified in Fig. 20-3 for (a) LPIM batch A and (b) HPIM batch G. The surface topographies of all batches are presented in Fig. 20-4. As listed in Table 20-2, the grain sizes of the LPIM specimens are in a range between 0.36 and 0.47 lm. The HPIM batch G has a corresponding grain size of 0.42 lm. The values of the porosity in Table 20-2 show significantly different values between 0.14 and 3.63% for the LPIM batches and a porosity of 0.37% for the HPIM batch. By comparing the porosities with the shear viscosities of the LPIM specimens in Table 20-1, the correlation between the two parameters becomes visible. Higher shear viscosities seem to result in higher porosities. Despite a comparable shear viscosity of batch A, the porosity

Fig. 20-3 Exemplary microstructures of (a) LPIM batch A and (b) HPIM batch G ZrO2 microbending specimens

20.4 Results and Discussion

Fig. 20-4 Surface topographies of the LPIM batches A–F and the micropowder injection molded batch G ZrO2 microbending specimens

of batch B is higher probably because of a worse degassing after LPIM. This can be seen by the obviously higher peak-to-valley roughness Rt = 3.25 lm for batch B compared with batch A with Rt = 1.05 lm. The surface topography of batch B shown in Fig. 20-4 reveals open semi-elliptical voids which are probably the negative image of air bubbles that must have been at the walls of the silicon negative

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20 Microstructure, Surface Topography and Mechanical Properties of Molded ZrO2 Microspecimens Table 20-2 Mean values of grain sizes, porosities and peak-to-valley-heights of batches A–G

Parameter

Grain size (lm) Porosity (%) Peak-to-valley Height Rt (lm)

Batch A

B

C

D

E

F

G

0.36 0.40 1.05

0.41 1.33 3.25

0.44 3.63 1.44

0.46 0.62 1.23

0.47 1.58 2.67

0.45 0.14 1.05

0.42 0.37 0.88

during the LPIM process. The surface roughnesses of batches C and D are Rt = 1.44 and 1.23 lm, respectively, and are comparable. The relatively low porosity of 0.62% of batch D in contrast to that of 3.63% of batch C can be related to the additional dwell time of 5 h at 180 8C during debinding, which was introduced to affect the edge geometry of the specimens. At this temperature, the thermal decomposition of the binder has not yet started and its viscosity is significantly reduced in comparison with the manufacturing temperature of 85 8C. It is argued that a low viscosity in combination with an additional dwell time is responsible for increasing the edge radii and for healing of defects such as pores. All specimen arrays of the LPIM batches were produced on a base plate. That of batch C was 2.5 times thicker compared with all other LPIM batches, which could have additionally influenced the debinding process. Batch E and batch F, which were nominally identically produced, have explicit differences in their porosities and surface qualities. Batch E has a porosity of 1.58% and a surface roughness of Rt = 2.67 lm, whereas batch F has lower values of 0.14% and Rt = 1.05 lm, respectively. The shear viscosities give no explanation in this case. The surface topography of batch E is characterized by many open pores, which lead to the higher Rt value. Furthermore, such a porosity is also found in the interior of the specimens of batch E. This means in correlation with batch B that there were also problems with the degassing of the LPIM feedstock, in this case probably starting already with the preparation and homogenization of the feedstock. Another conspicuous fact on comparison of batches E and F is the pronounced difference in the edge radii of the specimens which are listed for all batches in the following section. Batch E has a mean edge radius of 3.74 lm, which is relatively sharp compared with the 14 lm edge radius of batch F. This shows that there were obviously different surface tensions during debinding, which could be associated with an inhomogeneity of feedstock composition. In comparison with the LPIM specimens, the HPIM batch G has the smoothest surface, showing a roughness of Rt = 0.88 lm. Owing to the higher pressure of the HPIM process, the distinctly higher shear viscosity of the feedstock of batch G results, as already mentioned, in a relatively low porosity of 0.37%. Exemplary results of the residual stress analysis of batches C, D and G are presented in Table 20-3. All three batches show residual stresses between

20.4 Results and Discussion Table 20-3 Residual stress rRS of the low-pressure (batches C and D) and high-pressure (batch G) injection molded ZrO2 microbending specimens

Batch

rRS (MPa)

C D E

118 137 77

rRS = 77 and 137 MPa with a maximum scatter of DrRS = ± 90 MPa. Because of the technical limitations of the measurement in the x mode only a small x range (–32 £ w £ 328 and –28 £ w £ 288, respectively) could be used, which limits the accuracy of the results of the linear regression performed in the sin2w analysis. Therefore, the residual stresses measurements show relatively high standard deviations. However, the residual stresses determined are negligible compared with the bending strengths of all specimens investigated, which will be shown in the following section. 20.4.2 Determination of Mechanical Properties and Influencing Factors

Before giving the experimentally determined bending strengths of batches A–G, a short discussion of the strength values achievable by applying the Weibull size effect is presented. As already mentioned in Section 20.3, the characteristic bending strength of a specimen with a given size can be extrapolated to another size regarding the effective volumes [Equation (6 a)] or the effective surfaces [Equation (6 b)]. For such an extrapolation, macroscopic specimens, which have been produced by TOSOH using the same powder as for the LPIM microspecimens, are applied. These specimens with the dimensions 3 ´ 4 ´ 40 mm were tested with a three-point bending device which had a support distance of 30 mm. The mechanical tests with the macrospecimens resulted in a characteristic bending strength of r0,macro = 1200 MPa [24]. Owing to the lack of a Weibull

Table 20-4 Surface-to-volume ratio Seff/Veff of three-point bending macro- and microspecimens

Type of specimen

Weibull modulus, m

Effective surface area, S (mm2)

Effective volume, V (mm3)

Ratio Seff/Veff (mm–1)

Micro (0.2 ´ 0.2 ´ 1.2 mm); support distance 0.8 mm Macro (3 ´ 4 ´ 40 mm3); support distance 30 mm

9 10 9 10

0.018 0.016 12.9000 11.6529

0.000160 0.000132 1.8000 1.4876

110 120 7.17 7.83

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20 Microstructure, Surface Topography and Mechanical Properties of Molded ZrO2 Microspecimens Table 20-5 Weibull size effect by extrapolating the bending strength of standard macro specimens from TOSOH (3 ´ 4 ´ 40 mm) with r0,macro = 1200 MPa [24] to the presented microsize (0.2 ´ 0.2 ´ 1.2 mm), applying the effective volumes or the effective surfaces

Fictive Weibull modulus, mfictive

9 10

Extrapolated bending strength of microspecimens (MPa) Surface effect

Volume effect

2498 2322

3383 3050

modulus for the macrospecimens, the fictive Weibull moduli mfictive = 9 and 10 were chosen, which correlate well with standard Weibull moduli in the literature [25, 26] and to the m values from the microbending tests in the present work. The values of the effective surfaces and volumes of the macrospecimens are given in Table 20-4. By extrapolating the strength of the TOSOH specimens to the size of the microspecimens, whose effective surfaces and volumes are also listed in Table 20-4, the characteristic bending strengths for surface and volume effect given in Table 20-5 are derived. For mfictive = 9, the surface effect results in a characteristic bending strength of r0,surface = 2498 MPa and the volume effect in r0,volume = 3383 MPa. The slightly higher Weibull modulus of 10 leads to r0,surface = 2322 MPa for the surface effect and r0,volume = 3050 MPa for the volume effect. In addition to the obviously higher strength values of the microspecimens and the strong influence of the Weibull modulus on the calculation, it can be seen that the strengths of volume-defect dominated specimens are considerably higher than the values for surface-defect dominated specimens. The experimentally determined characteristic bending strengths r0 and the related Weibull moduli mcorr of batches A–G are shown in the Weibull diagram in Fig. 20-5. Generally, the results for the batches investigated can be approximated by straight Weibull lines. Only batch F shows pronounced deviations in the probability of failure up to 25%. The r0 values show a very large band width of 1417 up to 3181 MPa. The associated Weibull moduli which represent the scatter of single batches are in a usual range and show values between 7.2 and 14. In the following, the reasons for the large scatter of the r0 values will be analyzed by discussing the fracture surfaces in addition to the surface qualities, the edge radii and the porosities (see Table 20-2). Batch A has the highest characteristic bending strength of r0 = 3181 MPa and the lowest scatter of mcorr = 14 of the batches presented here. The related fracture surfaces in Fig. 20-6 show as main fracture origins slight unevennesses at the specimen sides loaded in tension. The specimens with higher strength values start to fail from pores smaller than 1 lm, which corresponds well with the low porosity of 0.40%. Very conspicuous are the large edge radii (14.46 lm) and the smooth specimen surfaces (Rt = 1.05 lm).

20.4 Results and Discussion

Fig. 20-5 Weibull diagram for the bending strengths rb and the related probabilities of failure, Pf, for batches A–F (LPIM) and G (HPIM)

Fig. 20-6 SEM images of fracture surfaces of batch A

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20 Microstructure, Surface Topography and Mechanical Properties of Molded ZrO2 Microspecimens

Batch B, which has r0 = 1914 MPa and mcorr = 7.2, fails mainly owing to semielliptical surface voids (Fig. 20-7), which lead to a mean surface roughness Rt = 3.25 lm, resulting in a strong notch effect. The depths of these voids range between 3 and 15 lm and the spreads are between 15 and 40 lm. As could be expected, the bending strengths are the lower the larger are the surface voids. Additionally, batch B has a relatively high porosity of 1.33%. The specimens with higher strength values fail again owing to volume defects, i.e. in this case pores beneath the surface of size 3–5 lm. The edge radii of 13.3 lm are comparable to those of batch A. As could be expected, batch C, which has the highest porosity of 3.63%, has a relatively low characteristic bending strength of r0 = 1568 MPa. The Weibull modulus of mcorr = 9.6 is in the middle of all the moduli presented. The fracture origins are mainly pores near or at the surfaces loaded in tension (Fig. 20-8). In individual cases the specimens fail owing to burrs originating from the micro-

Fig. 20-7 SEM images of fracture surfaces of batch B

Fig. 20-8 SEM images of fracture surfaces of batch C

20.4 Results and Discussion

milled mould. The edge radii of 6.16 lm are obviously sharper than those of batches A and B. The surface is of only moderate quality with a roughness Rt = 1.44 lm, which derives from numerous open surface pores on the scale of *3–5 lm in diameter. Batch D, which has r0 = 2279 MPa and mcorr = 11.3, fails similarly to batch C, mainly owing to pores located near or at the surface loaded in tension or at the edges (Fig. 20-9). However, the porosity of 0.62% is lower than that of batch C. Additionally, the pores have a diameter of *2–3 lm and are therefore distinctly smaller than the pores of batch C, which explains the higher strength level. Only single pores of specimens with lower bending strengths have a diameter of *5 lm. Compared with batch C, the smaller pores at the surfaces result in a slightly smaller roughness of Rt = 1.23 lm. The mean edge radius of 9.43 lm is slightly larger than the radius of batch C, probably being influenced by the longer debinding and the longer duration of stay of the binder (see Table 20-1). Batch E has a porosity of 1.58% and a surface roughness of 2.67 lm, which represent relatively bad structure and surface properties. Both values result in the lowest presented characteristic bending strength of r0 = 1417 MPa with a Weibull modulus of mcorr = 12.3. The fracture origins are exclusively open pores in the size range 3–8 lm or semi-elliptical surface voids with depths of *5 lm and spans between 10 and 20 lm (Fig. 20-10). The mean edge radius of batch E of 3.74 lm is the lowest for all the LPIM batches. Batch F has, in contrast to batch E, a porosity of 0.14% and a peak-to-valley roughness of 1.05 lm, which result in a distinctly higher characteristic bending strength of r0 = 2690 MPa. The Weibull modulus is mcorr = 9.2. The specimens of batch F fail similarly to batch E mainly owing to pores located in or at the surfaces loaded in tension. However, the pores which have a diameter of *2–3 lm are smaller (Fig. 20-11). Single specimens of this batch fail also owing to semielliptical surface voids with dimensions of 3–5 lm, which also are smaller. The edge radii of *14 lm are, as already mentioned in the preceding section, significantly larger than those of batch E.

Fig. 20-9 SEM images of fracture surfaces of batch D

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20 Microstructure, Surface Topography and Mechanical Properties of Molded ZrO2 Microspecimens

Fig. 20-10 SEM images of fracture surfaces of batch E

Fig. 20-11 SEM images of fracture surfaces of batch F

In comparison with the LPIM specimens, the HPIM batch G has a relatively low porosity of 0.37% and a very low surface roughness of 0.88 lm, but a low characteristic bending strength of r0 = 1789 MPa. The scatter of the bending strengths represented by mcorr = 7.9 is slightly higher than for most of the other batches. The HPIM specimens fail mainly owing to open and enclosed pores of size 4–8 lm located at the edges which have a mean radius of 2.63 lm (Fig. 2012). Very often several pores can be detected in the area of the fracture origin. Specimens with higher bending strengths fail owing to smaller single pores of size 2–3 lm located beneath the surfaces loaded in tension (Fig. 20-12, right). In order to correlate the microstructural, geometric and mechanical properties of batches A–G, the characteristic bending strength is presented in Figs. 20-13 and 20-14 versus the surface roughness or porosity and the edge radius. Despite a relatively similar surface roughness of batches A, D, F and G with values of Rt between 0.88 and 1.23 lm, the bending strengths differ, varying between 1789 and 3181 MPa. By additionally correlating the bending strength and the surface roughness with the edge radius, the bending strength increases with increasing

20.4 Results and Discussion

Fig. 20-12 SEM images of fracture surfaces of batch G

edge radius. In contrast, batches with approximately identical edge radii, e.g. batchs B and F (radius *14 lm) or batches E and G (radius *3 lm), show expectedly a decreasing strength with increasing surface roughness. In comparison with batch F, batch A has a higher bending strength, probably due to its smaller grain size [27] of 0.36 lm compared with 0.45 lm for batch F. By exchanging the surface roughness with the porosity in Fig. 20-14, the correlation looks relatively similar to Fig. 20-13. Batches A, D, F and G, which show an increasing strength with increasing edge radius, all have relatively low porosities with values between 0.14 and 0.62%. As could be expected from experiments with macroscopic brittle polycrystalline specimens [27] and seen here with batches A, F and B or batches E and G, increasing porosity and almost constant edge radii lead to decreasing strength values. The slightly higher characteristic bending strength of batch C of 1568 MPa and a porosity of 3.63% in contrast to batch E with 1417 MPa and 1.58% can be explained by the difference

Fig. 20-13 Characteristic bending strength versus surface roughness and edge radius

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20 Microstructure, Surface Topography and Mechanical Properties of Molded ZrO2 Microspecimens

Fig. 20-14 Characteristic bending strength versus porosity and edge radius

in pore size. As can be seen in Figs. 20-8 and 20-10, by comparing the fracture surfaces of batches C and E, the pores of batch E which cause failure are nearly two times larger than the critical pores of batch C. The larger pores located in or beneath the specimen surface loaded in tension result in a stronger notch effect and therefore reduce additionally the bending strength. According to these correlations, it can be concluded that the highest bending strengths are reached exclusively with specimens which have a low porosity up to 0.5%, a smooth surface with an Rt of about 1 lm and a large edge radius of at least 9 lm. This fits accurately with batches A and F. Most of these specimens tend to fail owing to volume defects, i.e. small pores of size 2 lm. Considering the bending strength level, the next two batches are D and B. Batch D has a slightly higher porosity and a smaller edge radius. Batch B has an edge radius that is comparable to those of batches A and F, but the surface roughness and the porosity are distinctly higher. The HPIM batch G, which has a slightly lower bending strength than batch B, has very smooth surfaces and a low porosity. However, the few pores are located very unfavorably in the relatively sharp edges so that nearly all specimens start to fail from the pores at the edges. Finally, batches C and E, which have the lowest strength values, show higher Rt and porosity values combined with smaller edge radii. The specimens of the batches with higher surface roughness, i.e. batches B, C and E, start to fail mostly from open pores at the surfaces loaded in tension, which shows that the failure behavior is mainly dominated by surface defects. The very high influence of the surface quality can also be seen by comparing the surface to volume ratios of macro- and microspecimens in Table 20-4: the ratio for the microspecimens is about 15 times larger than that for the macrospecimens. However, if the failure origins are beneath the surface loaded in ten-

20.4 Results and Discussion

sion, the bending strengths increase and the Weibull size effect regarding the effective volumes becomes valid. 20.4.3 Simulation of Influencing Factors on the Bending Strength

In experiments and practical applications, there is always a superposition of several factors such as edge radius, void geometry and roughness influencing the bending strength which cannot be divided or weighted easily. However, in order to design and dimension microparts made of brittle ceramics such as ZrO2, it is essential to estimate the effect of single influencing factors. Therefore, FEM calculations were performed by simulating bending specimens with a sharp 908 edge or an edge with a 20 lm radius. Owing to the observation that most of the specimens failed starting from surface defects, i.e. open pores or semi-elliptical surface voids, a single pore located in the middle or at the edge of the surface loaded in tension was also included in the model. The pore size was varied as 3, 6, 15 and 30 lm. As already described in Section 20.4.1, the grain sizes of all investigated batches are smaller than 0.5 lm. This results in combination with a cross-sectional geometry of 200 ´ 200 lm and the assumption of statistically distributed grain orientations in linear elastic and quasi-isotropic material behavior. Owing to a slightly lower section modulus of the specimen with a 20 lm edge radius resulting from less material at the edges, the stress values of both edge geometries could not be correlated directly. Therefore, the results of the FEM simulations are presented as relative stress values calculated by correlating the stress values along the section lines given in Fig. 20-15 with the local stress at z = 0 on the surface loaded in tension of the corresponding model without pores. The relative stress values along the section lines are shown in Fig. 20-16. The maximum values of these relative stresses are given for each pore geometry in Table 20-6. As indicated in Fig. 20-15, the points of maximum stress are normally not located directly on the surface loaded in tension but in the interior next to the pores. In comparison with the relative stress values of

Fig. 20-15 Results of FEM simulations – analyzed section lines and points of maximum stress relating to the pore location; exemplary diameter of pores is 6 lm

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20 Microstructure, Surface Topography and Mechanical Properties of Molded ZrO2 Microspecimens

Fig. 20-16 Results of FEM simulations – relative stress comparison between r(z) and r(z = 0) of the model without pores at the line of maximum deflection

Table 20-6 Results of FEM simulations – additional stresses (%) due to pores located at z = 0 or at the edge considering the maximum points of stress

Pore diameter (lm)

3 6 15 30

20 lm edge radius

Sharp 908 edge Pore at z = 0

Pore at edge

Pore at z = 0

Pore at edge

+35 +60 +80 +86

+29 +55 +75 +81

+37 +59 +80 +87

+38 +57 +90 +122

20.4 Results and Discussion

the section lines in Fig. 20-16, the maximum relative stress values occurring besides the section lines do not differ by more than 5%. As could be expected, along the section lines and at the point of the maximum relative stress, the models with both edge geometries generally show increasing relative stress values with increasing pore size. By comparing the relative stresses of each pore size for both the sharp and rounded edges, the simulations with a single pore in the middle of the surface loaded in tension show, as expected, comparable paths of the section lines and maximum values. In contrast, the different sized pores in the 20 lm radius show distinctly higher relative stress values than the pores in the sharp 908 edge. Generally, the difference in the relative stress values between both edge geometries increases with increasing pore size: However, the difference between the relative stress maxima of the 3 lm pores amounts to 9%, which is obviously higher than the 2% difference between the 6 lm pores. This can be traced back to the relation between pore size and mesh size of the FEM models. The smallest elements of the FEM models in the surrounding of the pores have edge lengths of about 0.5 lm, which means that the 3 lm pore is modeled obviously rougher than the 6 lm pore, which may result in some inaccuracy in the stress calculation. The difference between the maximum relative stresses of the 15 lm pore amounts to 15% and between the 30 lm pores to 41%. These results seem to show in the first instance the opposite of the observed edge effect in the experiments, but by taking into account the origins of failure of the batches investigated it becomes clear that most of the specimens, especially those with the highest bending strengths, fail owing to pores or semi-elliptical voids located at the surfaces loaded in tension. The specimens with rounded edges (radius > 9 lm) usually have no pores directly at the edge surfaces, which could be explained by the surface tension that leads to the radius and additionally to a smooth surface. The few specimens which failed starting from defects in rounded edges show in the case of higher strength values pore sizes of only 1–2 lm, e.g. the batch D specimen in Fig. 20-9 with rb = 2449 MPa or the batch F specimen in Fig. 20-11 with rb = 2057 MPa. Distinctly lower strength values are reached with rounded specimen edges only in individual cases, e.g. the batch D specimen in Fig. 20-9 with rb = 1519 MPa and a pore size of about 6 lm. In order to validate the calculated maximum relative stress factors as a kind of ‘notch factor,’ they could be applied to the strength values of single specimens which failed owing to pores with similar locations and/or sizes. By assuming that these specimens have no other defects in the size of the detected pores from which failure started and that other existing defects are distinctly smaller, the bending strengths can be adjusted to higher values by applying the FEM factors. The thus calculated higher strength values can be compared with the highest strength values of batches A and F which correspond to the smallest defect sizes. This procedure is only valid for batches with low porosities. Therefore, batches C and E, both of which have a porosity higher than 1.5%, were not taken into account. Additionally, the calculated strength values can be compared with the maximum values of the Weibull size effect.

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The batch G specimens show a nearly sharp 908 edge, having an edge radius of 2.63 lm. The failure origins of this batch are mainly located in these sharp edges in combination with pores of size between 4 and 8 lm. For the comparison, the specimen with rb = 1838 MPa in Fig. 20-12 is chosen. The size of the pore accumulation in the edge could be assumed to be 6 lm, which results corresponding to Table 20-6 in an additional relative stress of 55%. By multiplying the relative stress with the bending strength of the specimen, a fictitious strength of 2849 MPa is reached. This is relatively close to the results for batches A and F, which show the highest bending strengths owing to their relatively low defect population, and also relatively close to the strength values estimated by the Weibull size effect regarding the effective volumes: 3050 MPa for mfictive = 10 or 3383 MPa for mfictive = 9. In contrast to batch G, the specimens of batch B have a relatively large edge radius (13.3 lm) and fail mainly owing to semi-elliptical voids at the surface loaded in tension. By multiplying the bending strength 1945 MPa of the specimen in Fig. 20-7 that failed owing to a void of size about 5 lm deep and 15 lm wide with the roughly corresponding stress factor of 59%, an estimated stress of 3093 MPa is reached. This value also corresponds relatively well with the Weibull size effect values and with the high strength values of batches A and F. In order to give an example for specimens with relatively small pores as the origin of failure, specimens of batch D are used. The batch D specimens fail mainly owing to pores with a diameter of about 1–3 lm located at the surface loaded in tension. For the calculation, the specimen in Fig. 20-9 with rb = 2449 MPa is chosen. This specimen failed owing to an open pore with a diameter of about 1–2 lm located at the edge with a radius of *9.43 lm. The corresponding stress factor in Table 20-6 is 38%, which results in an estimated strength of 3380 MPa and again fits very well with the Weibull size effect and the high strength values of batches A and F. Generally, the validation shows that the stress factors calculated by FEM simulation show good agreement with the experiments even if the simulations could only give a rough estimation factor owing to the strong simplification and the limited number of simulated pore sizes and pore shapes.

20.5 Conclusions and Outlook

The microstructures, residual stresses, surface topographies and bending strengths of microspecimens made of tetragonal polycrystalline ZrO2 partially stabilized with 3 mol% Y2O3 and produced by low-pressure (LPIM) or high-pressure injection molding (HPIM) have been presented. Additionally, the influencing factors on the bending strength, i.e. porosity, surface topography and edge geometry, were determined and partly simulated by finite element method (FEM) calculations. The microstructures of the investigated batches show homogeneous grain structures with mean grain sizes between 0.36 and 0.47 lm. The porosities are

20.5 Conclusions and Outlook

between 0.14 and 3.63% for the LPIM batches and 0.37% for the HPIM batch. The peak-to-valley roughness of the HPIM batch amounts to 0.88 lm and those of the LPIM specimens are between 1.05 and 3.25 lm. Exemplarily, the residual stresses of two LPIM batches and of the HPIM batch were shown. The residual stresses are close to zero and therefore could be neglected with respect to the bending strengths. The characteristic bending strength of all batches varies in a range between 1417 and 3181 MPa for the LPIM specimens and amounts to 1789 MPa for the HPIM batch. The wide range of the mean edge radii of the LPIM specimens between 3.74 and 14.46 lm is very conspicuous and could be explained by different debinding processing (in relation to feedstock composition; see also Chapter 12). The mean edge radius of the HPIM batch is 2.63 lm, which shows in combination with the surface roughness the very high contour accuracy of the process. As could be expected, the analysis of the fracture surfaces shows a strong influence of the surface roughness and of the locations and sizes of critical pores on the bending strength. However, the correlation between bending strength, surface roughness or porosity and the edge radii of the specimens shows that the bending strengths increase not only with smoother surfaces and lower porosity but also with increasing edge radius. The characteristic bending strengths of the two LPIM batches which have porosities below 0.5%, a peak-to-valley-roughness of 1.05 lm and edge radii of about 14 lm fit very well with the Weibull size effect considering the effective volumes. FEM simulations of the bending tests with specimens that have a sharp 908 edge or an edge with a 20 lm radius, both combined with a single pore in the middle or at the edge of the surface loaded in tension, result in stress factors which could be validated by applying them to batches with porosities lower than 1.5% and by comparing them with the two high-strength batches with the smallest defects. Further experiments will be performed in order to improve and broaden the correlation between characteristic bending strength, surface roughness, porosity and edge radius. Additional FEM simulations with enclosed pores and crooked specimen alignment on the bending device so that one edge on the surface loaded in tension is more stressed than the other will also help by establishing more reliable material data to be used within a dimensioning concept for microparts made of brittle engineering ceramics.

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20.6 References 1 J. Marz, N. Burkardt, A. Albers, in: Conf. 2

3 4

5 6 7

8

9

10

11 12

13

14

15

Proc. EUSPEN; 2003, Vol. 1, pp. 23–26. D. Spath, T. Barrho, M. Knoll, D. Löhe, V. Schulze, wt Werkstattstechnik online 2003, 93, H.3, pp. 136–140. www.sfb499.de W. Bauer, G. Bartelt, H.-J. RitzhauptKleissl, J. Hausselt, in: Conf. Proc. Key Engineering Materials; 2001, Vol. 206–213, pp. 329–332. W. Bauer, R. Knitter, J. Mater. Sci. 2002, Vol. 37, 3127–3140. R. Knitter, W. Bauer, D. Göhring, J. Mech. Eng. Sci. 2003, 217, No. 1, 41–51. H.-J. Ritzhaupt-Kleissl, R. Ruprecht, W. Bauer, V. Piotter, J. Hausselt, in: Proc. Pulvermetallurgie in Wissenschaft und Praxis, Vol. 15; Düsseldorf: VDI Werkstofftechnik, 1999, pp. 185–204. L. Merz, S. Rath, V. Piotter, R. Ruprecht, J. Ritzhaupt-Kleissl, J. Hausselt, Microsyst. Technol. 2002, 8, pp. 129–132. L. Merz, S. Rath, V. Piotter, R. Ruprecht, J. Hausselt, Microsyst. Technol. 2004, 10, pp. 202–204. V. Piotter, G. Oerlygsson, R. Ruprecht, J. Hausselt, in: Conf. Proc. 4th EUSPEN; 2004, pp. 5–6. G. Schneider, L. Postlter, L. Wührl, Structure 1996, 29, pp. 19–23. B. Kagerer, R. Brodmann, J. Valentin, J. Filzek, U. Popp, Proc. SPIE 2002, 4773, pp, 52–62. Geometrische Produktspezifikationen, Oberflächenbeschaffenheit, Tastschnittverfahren, Regeln und Verfahren für die Beurteilung der Oberflächenbeschaffenheit; DIN EN ISO 4288, Berlin: Beuth Verlag, 1998. T. Wroblewski, O. Clauss, H.-A. Crostack, A. Ertel, F. Fandrich, Ch. Genzel, K. Hradil, W. Ternes, E. Woldt, Nucl. Instrum. Methods Phys. Res. A 1999, 428, pp. 570–582. T. Wroblewski, in: 6th European Conference on Residual Stresses – ECRS 6, Coim-

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23 24

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bra, Portugal, Materials Science Forum; 2002, Vols. 404–407, pp. 121–126. E. Macherauch, B. Eigenmann, ‘Röntgenographische Untersuchung von Spannungszuständen in Werkstoffen,’ Mater.-wiss. u. Werkstofftech. Teil I & II 26, Teil III & IV 27, 1995–1996. K. Obergfell, T. Beck, V. Schulze, D. Löhe, Materialprüfung 2000, Vol. 42 (10), pp. 391–395. M. Auhorn, T. Beck, V. Schulze, D. Löhe, Microsyst. Technol. 2002, 8, 109–112. M. Auhorn, B. Kasanická, T. Beck, V. Schulze, D. Löhe, Z. Metallkd. 2003, 94, pp. 599–605. D. Gronych, M. Auhorn, T. Beck, V. Schulze, D. Löhe, Z. Metallkd. 2004, 95, pp. 551–558. D. Munz, T. Fett, Ceramics, Mechanical Properties, Failure Behaviour, Materials Selection, 2nd edn.; Berlin: Springer-Verlag, 2001. Prüfung von keramischen Hochleistungswerkstoffen, 4-Punkt-Biegeversuch, Statistische Auswertung, Ermittlung der WeibullParameter; DIN 51 110 Teil 3, Berlin: Beuth Verlag, 1993. G. D. Quinn, J. Am. Ceram. Soc. 2003, 86, pp. 508–510. Tosoh Zirconia Powder E-Grades, Data Sheet of the Advanced Ceramics Department, Tokyo: TOSOH, 1999. H. Salmang, H. Scholze, Keramik Teil 1 – Allgemeine Grundlagen und wichtige Eigenschaften, 6th edn; Berlin: SpringerVerlag, 1982. R. Hessert, Bearbeitungseigenspannungen, Randschichtschädigungen und Festigkeiten geschliffener Al2O3- und ZrO2-Keramiken; Dissertation, Universität Karlsruhe (TH), 1998. F. P. Knudsen, J. Am. Ceram. Soc. 1959, 42, pp. 376–387.

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21 Tribological Characterization of Mold Inserts and Materials for Microcomponents J. Schneider 1), K.-H. Zum Gahr 1, 2), J. Herz 1) 1) Institute of Material Science and Engineering II, University of Karlsruhe (TH), Germany 2) Institute for Material Research I (IMF I), Research Center Karlsruhe, Germany Abstract The tribological performance of microstructured metallic specimens simulating mold inserts and self-mated alumina and zirconia sliding pairs was characterized by using laboratory tribometers. The wear behavior of mold insert specimens made from unalloyed, low- and high-alloyed steels, electroplated nickel and brass CuZn37 was studied using a specially developed laboratory tribometer simulating micropowder injection molding (lPIM) with zirconia and alumina feedstocks. The highest wear resistance was measured for the high-alloyed steel X175VCrMo9-5, which was produced by a powder metallurgy process. This steel showed a homogeneous microstructure consisting of a martensitic matrix and about 20 vol.% fine carbides. The wear resistance of single-phase electroplated nickel was significantly higher than that of the heat-treated tool steels 30CrMo6 and X38CrMoV5-1 of higher hardness. These results indicate that wear behavior in lPIM depends more on microstructural parameters such as homogeneity than on hardness of the materials. Wear behavior was also strongly influenced by the characteristics of the feedstock and the test parameters, such as the temperature used. Friction and wear of self-mated zirconia and alumina microparts produced by lPIM were characterized using a laboratory tribometer (system: ball-on-disc) in unidirectional sliding contact. The effect of normal load was studied in ambient air with a relative humidity of 50% and under water lubrication, respectively. The results proved that both zirconia and alumina can be favorably used for tribologically high-loaded micromechanical components particularly under water lubrication. Alumina showed lower friction but higher wear, whereas zirconia yielded both reduced friction and wear on comparing the water-lubricated with the unlubricated tests.

Keywords microtribology; wear; friction; microtechnology; powder injection molding; ceramic feedstock; mold wear; erosion; abrasion; water lubrication

Advanced Micro and Nanosystems Vol. 4. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31493-8

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21.1 21.2 21.2.1 21.2.1.1 21.2.1.2 21.2.2 21.2.2.1 21.2.2.2 21.3 21.3.1 21.3.1.1 21.3.1.2 21.3.2 21.3.2.1 21.3.2.2 21.4 21.4.1 21.4.2 21.5 21.6 21.7

Introduction 580 Materials and Experimental Methods Materials 582 Mold Insert Specimens 582

582

Specimens for Microtribological Sliding Tests 583 Tribological Tests 584 Micropowder Injection Molding Tests 584 Microtribological Sliding Tests 585 Results 586

Microstructure 586 Mold Insert Specimens 586 Specimens for Microtribological Sliding Tests 587 Tribological Properties 588 Wear of Mold Insert Specimens 588 Behavior in Microtribological Sliding Tests 591 Discussion

596

Wear of Mold Inserts in Micropowder Injection Molding Test 596 Friction and Wear During Microtribological Sliding Tests 599 Conclusions 600 Acknowledgments References 601

600

21.1 Introduction

The success of micromechanical devices depends both on the utilization of costeffective manufacturing technologies and on the utilization of a broad range of polymeric, metallic and ceramic materials [1, 2]. Micropowder injection molding (lPIM) is considered to be a very promising technology for the large-scale production of microcomponents, which can meet these requirements [3–6] (see Chapter 10). Nowadays, mold inserts made from etched silicon, electroplated nickel or micromachined brass are typically used in injection molding of polymeric microcomponents (see also Chapters 8 and 10). However, with increasing mechanical and tribological demands in micromechanical systems, e.g. microturbines or planetary gear boxes with high output torques, it is necessary to develop suitable metallic or ceramic microcomponents offering higher critical loads than typical polymeric components. Several studies on micropowder injection molding with ceramic or metallic feedstocks showed that the development of mold inserts with increased wear resistance compared with those used in microplastics processing is essentially required [7–11] (see Chapters 11 and 12). Promising wear-resistant materials should be, for example, those which are nowadays successfully used for macroscopic mold inserts [12, 13]. As the heat treatment of a completely patterned micromold insert made from steel would typically lead to distortion or even destruction of the microcav-

21.1 Introduction

ities, the mold inserts have to be machined after heat treatment. Micromilling, drilling or grinding, laser ablation or electrodischarge manufacturing (EDM) are techniques which have been successfully used to fabricate mold inserts from various heat-treated steels and cemented carbides [14–17] (see also Chapters 5, 6 and 7). Wear-resistant coatings such as hard chrome, electroless nickel, CrN or TiN, which are successfully used in macroscopic plastics processing, are not simply applicable for microscopic mold inserts owing to the typically small tolerances of cavity geometries and problems in ensuring homogeneous coatings within small cavities of high aspect ratio [18, 19]. Whereas there have been only a small number of studies on the specific tribological problems of mold inserts used in lPIM [20–22], many studies have characterized the tribological behavior of macroscopic mold inserts and other machine parts in plastics processing and powder injection molding [12, 13, 18, 19, 23–25]. In addition to expensive in situ experiments using injection molding machines, several simplified laboratory tribometers have been developed [13, 23, 24]. Studies with filled and unfilled feedstocks showed that both abrasive and erosive wear occur depending on the process parameters, the rheological properties of the molten feedstock, the material and the geometry of the mold insert cavities. Aggressive products (e.g. HCl) from organic binders or flame retardants which are set free from the molten feedstock can additionally lead to corrosion, especially at higher temperatures. The performance of micromechanical systems with movable components is strongly affected by the tribological behavior of the materials used. High friction forces reduce the usable output torque (kinetic friction) or even prevent (static friction) the operation of microturbines, motors and gears. Even very small amounts of wear resulting in loose wear debris can cause the loss of function of micromechanical devices. Hence friction and wear have to be recognized as very important issues in microtechnology. Several studies on the friction and wear of tribologically loaded microcomponents clarified that results from tribological tests on the macroscale cannot be simply transferred to the microscale. Because of the large surface area-to-volume ratio in micromechanical systems, adhesion and/or capillary forces become critical and can even surpass externally applied forces. Therefore, the friction and wear behavior of any material pair which is considered for utilization in microtribological systems has to be systematically characterized on the microscale [26–32]. In addition to experimental studies on the microscale, finite element simulations can be helpful for a better understanding of critical loading conditions in the tribological contact zone (see also Chapter 22). Ceramic materials such as zirconia and alumina offer some major advantages over polymeric and metallic materials owing to their unique combination of low density, high strength, stiffness and hardness, temperature and chemical resistance and low adhesion tendency. Therefore, ceramic materials should be applicable both under high contact stresses and sliding speeds and in aggressive environments. Macroscopic tests with self-mated zirconia and alumina pairs showed that friction and wear depend strongly on the operating (e.g. load,

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speed) and environmental (e.g. temperature, humidity, liquid media) conditions [33–37]. The tribological behavior of alumina pairs is characterized by a transition from mild to severe wear, owing to a change from predominantly tribochemical reaction to surface damage by microfracture. During tests in humid air or under lubrication with aqueous solutions, the formation of soft aluminum hydroxide films with low shear strength in the tribological contact zone can effectively reduce friction in the mild wear regime [38–42]. Tribological tests with self-mated zirconia showed that in both unlubricated and water-lubricated sliding contact, relatively high friction and wear can occur [39, 43–45]. Depending on the operating parameters, the tribological behavior of zirconia ceramics partially stabilized with yttria (e.g. 3Y-TZP: ZrO2 with 3 mol% Y2O3) can be affected by the stress-induced transformation from the tetragonal to the monoclinic phase and hydrothermal destabilization of the metastable tetragonal phase by water due to the leaching of Y2O3 and the formation of Y(OH)3 [46–48]. The aim of this study was to characterize the tribological performance of micromachined mold inserts made from various unalloyed, low- and high-alloyed steels and alumina and zirconia sliding pairs using laboratory tribometers. Experiments were carried out using a specially developed tribometer simulating lPIM with zirconia and alumina feedstock at 170 and 190 8C, respectively. Influences of the microstructures of the steels were studied and compared with results achieved from experiments with electroplated nickel and brass CuZn37. Furthermore, the microtribological performance of self-mated zirconia and alumina ceramics produced by lPIM was characterized in unidirectional sliding contact. Operating parameters for these tests were chosen on the basis of a system analysis of the microturbine and gear used as demonstrators in the SFB 499.

21.2 Materials and Experimental Methods 21.2.1 Materials 21.2.1.1 Mold Insert Specimens Tribological tests simulating the wear behavior of mold inserts in lPIM were carried out using steel specimens. Heat treatment of the specimens was done in close collaboration with the Institute of Materials Science and Engineering I (see also Chapter 9). The unalloyed plain carbon steel C 80 (0.8% C) was tested with normalized microstructure (austenitization at 830 8C in vacuum for 30 min and cooling within the furnace). Furthermore, normalized specimens were tempered at 700 8C for 240 min and specimens austenitized at 830 8C in vacuum for 30 min were oil quenched and tempered at 480 8C for 120 min, respectively. The low-alloyed steel 30CrMo6 (Toolox 44, SSAB Oxelösund, Sweden; 0.3% C, 1.35% Cr) was hardened and tempered at 590 8C by the manufacturer. According to the manufacturer, the tool steel Toolox 44 showed high purity and homo-

21.2 Materials and Experimental Methods

geneity comparable to electro slag remelted steels. The high-alloyed vacuum remelted tool steel X38CrMoV5-1 (0.38% C, 5% Cr, 1% Mo; Böhler, Germany) was austenitized in air at 990 8C for 20 min, quenched in oil and tempered at 560, 520 and 610 8C (2 ´ 60 min). Specimens of X38CrMoV5-1 with a bainitic microstructure were obtained by an isothermal transformation at 350 8C. The high-alloyed steel X175VCrMo9-5 (CPM9V, Crucible, USA; 1.75% C, 9% V, 5% Cr), produced in a powder metallurgy process, was austenitized in vacuum at 1120 8C for 60 min, oil quenched and tempered at 560 8C (2 ´ 60 min). Specimens made from electroplated nickel and brass CuZn37 were chosen as reference. Surface finishing of the specimens by grinding and micromilling was done at the Institute of Production Science (see also Chapter 5) and resulted in an average surface roughness Ra (center line average, CLA) between 0.3 and 0.36 lm.

21.2.1.2 Specimens for Microtribological Sliding Tests The materials used for the microtribological sliding tests are listed in Table 21-1. Discs of zirconia and alumina with a diameter of 8 mm were microinjection molded at the Forschungszentrum Karlsruhe, Institute for Materials Research III (see also Chapter 12) using commercial feedstocks (ZrO2, CATAMOLD TZPA; Al2O3, CATAMOLD AOF; BASF, Germany). These feedstocks were based on a POM (polyoxymethylene) binder with about 50 vol.% of zirconia (ZrO2, CATAMOLD TZP-A, Fig. 21-1 c) or 58 vol.% alumina (Al2O3, CATAMOLD AOF, Fig. 21-1 d) and some other minor additions. The size of the fine ceramic particles ranged from 0.3 to 0.4 lm for ZrO2 and from 0.4 to 0.6 lm for Al2O3. Whereas the zirconia particles were mostly of spherical shape, the alumina particles were partially sharp-edged [49]. Balls made of commercial TZP-ZrO2 with 3 mol% of Y2O3 and of alumina (both from SWIP, Switzerland) with a diameter of 1.588 mm were used as counter bodies.

Table 21-1 Properties of ZrO2 and Al2O3 ceramics used for the microtribological sliding tests (* = specifications of the manufacturer)

Parameter

Density (103 kg/m3) Mean grain size (lm) Young’s modulus (GPa) Hardness (HV500)

Discs (lPIM)

Balls

ZrO2

Al2O3

ZrO2

Al2O3

6 0.5 200 1345

3.74 0.7 393 2183

6.05 1 200 1236

3.97 10 380 2100

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21 Tribological Characterization of Mold Inserts and Materials for Microcomponents

Fig. 21-1 Schematic description of (a) the tribosystem simulating injection molding, (b) the mold insert specimens and scanning electron micrographs of the (c) ZrO2 and (d) Al2O3 feedstocks

21.2.2 Tribological Tests 21.2.2.1 Micropowder Injection Molding Tests Wear tests were carried out using a specially developed laboratory tribometer (Fig. 21-1 a) simulating microinjection molding [21] with the two different CATAMOLD feedstocks (BASF, Germany), which were also used for the fabrication of the microtribological disc specimens. A piston with two pairs of micromachined specimens (Fig. 21-1b) was moved up and down within a cylinder filled with a volume of 14 ± 1 cm3 of feedstock. During each stroke of the piston with an average velocity of 2 mm/s, the molten polymeric part of the feedstock was pushed through a square slot (1.5 ´ 1.5 mm) with a pressure between 0.6 and 1.1 GPa at an average flow velocity of 245 mm/s. Before starting the tests, the complete tribometer was heated to 170 or 190 8C. These operating parameters were chosen based on an analysis of the process in cooperation with the Institute for Materials Research III at the Forschungszentrum Karlsruhe (see also Chapter 12). Depending on the number of simulated injections, wear-induced mass loss was measured using an electronic balance with a scale of 10–5 g. The reported values of the volumetric wear WV which were calculated from the mass

21.2 Materials and Experimental Methods

loss Wm according to WV = Wm/q are the averages of at least two individual tests under the same test conditions. Tribological testing was accompanied by analysis of the worn surfaces using scanning electron microscopy and optical profilometry.

21.2.2.2 Microtribological Sliding Tests Unidirectional sliding tests were carried out in a laboratory ball-on-disc microtribometer developed at the Institute of Materials Science and Engineering II. The surfaces of the disc specimens were ground to average surface roughness values of Ra = 0.12 lm (ZrO2) and 0.36 lm (Al2O3) using a diamond wheel of 15 lm grit size. The ZrO2 and Al2O3 balls used as counter bodies were delivered by the manufacturer in a polished condition with a surface roughness Ra = 0.12 lm. The microtribological tests were run over a sliding distance of s = 500 m in ambient air with a relative humidity of 50% and in distilled water at normal loads of FN = 200, 400 and 800 mN and a sliding speed of v = 400 mm/s (Fig. 21-2 b). The parameters for the laboratory tests were chosen on the basis of a system analysis of the microturbine and planetary gear box used as demonstrators (Fig. 21-2 a), which was done in cooperation with the Institute of Product Development (see also Chapters 1 and 2). Normal force FN and friction force FR were continously measured during all tests. The amount

; distilled water

Fig. 21-2 Schematic description (a) of the demonstrator system (microturbine with planetary gear box) and (b) the laboratory tribosystem (ball-on-disc) for unidirectional sliding, operating conditions and geometries of the specimens

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21 Tribological Characterization of Mold Inserts and Materials for Microcomponents

of wear was determined after the tests by measuring the contact diameter on the worn ball specimens using an optical microscope and by optical and tactile surface profilometry of the disc specimens. Furthermore, the worn surfaces were examined using scanning electron microscopy.

21.3 Results 21.3.1 Microstructure 21.3.1.1 Mold Insert Specimens Fig. 21-3 shows scanning electron micrographs of the materials used for the mold insert specimens. Compared with the normalized steel C 80 with a fine pearlitic microstructure (Fig. 21-3 a), the steel C 80 tempered at 700 8C for 240 min showed a coarser pearlite with partially coalesced and spheroid cemen-

Fig. 21-3 Scanning electron micrographs of the microstructures of the materials used as mold insert specimens: (a) steel C 80 normalized; (b) steel C 80 normalized and tempered at 700 8C; (c) steel X38CrMoV5-1 hardened and tempered at 560 8C; (d) PM steel X175VCrMo9-5; (e) electroplated nickel; (f) brass CuZn37

21.3 Results

tite particles (Fig. 21-3 b). The high-alloyed steel X38CrMoV5-1 had a martensitic microstructure with less than 5 vol.% carbides up to a diameter of 1.5 lm (Fig. 21-3 c). The microstructure of the PM steel X175VCrMo9-5 was characterized by about 20 vol.% fine, homogeneously distributed carbides within the martensitic matrix (Fig. 21-3 d). The nickel (Fig. 21-3 e) showed a fine, columnar grain structure typical for the electroplating process used. Single-phase a-brass CuZn37 had a bimodal grain size distribution and a much coarser microstructure than all other materials (Fig. 21-3 f).

21.3.1.2 Specimens for Microtribological Sliding Tests Fig. 21-4 shows scanning electron micrographs of the materials used in the microtribological tests. The mean grain size was 0.5 lm for the microinjection molded ZrO2 used for the disc specimens (Fig. 21-4 a) and 1 lm for the ZrO2 balls (Fig. 21-4 c). Both materials consisted of the tetragonal ZrO2 phase owing to the partial stabilization with 3 mol% Y2O3. The zirconia of the disc specimens contained about 10 vol.% Al2O3 (dark phase in Fig. 21-4 a) and 90 vol.% ZrO2 (light phase in Fig. 21-4 a). The microstructure of the Al2O3 balls (Fig. 214 d) with a mean grain size of about 10 lm was much coarser than that of the alumina disc specimens of 0.7 lm (Fig. 21-4 b). On the other hand, the microstructure of the lPIM alumina disc specimens showed a higher porosity of about 5 vol.% (Fig. 21-4 b).

Fig. 21-4 Scanning electron micrographs of the microstructure of the (a) ZrO2 (lPIM, 10 vol.% Al2O3) and (b) Al2O3 (lPIM) disc specimens, (c) the ZrO2 (3 mol% Y2O3) and (d) the Al2O3 balls used for the microtribological sliding tests

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21 Tribological Characterization of Mold Inserts and Materials for Microcomponents

21.3.2 Tribological Properties 21.3.2.1 Wear of Mold Insert Specimens Fig. 21-5 shows the volumetric wear of the steel X38CrMoV5-1 (hardened and tempered at 560 8C) as a function of the number of simulated injections with zirconia feedstock at 170 and 190 8C. Volumetric wear calculated from the mass loss continuously increased with increasing number of strokes (simulated injections). The amount of wear was significantly greater at a feedstock temperature of 170 8C than at 190 8C. Scanning electron micrographs of worn surfaces of the steel X38CrMoV5-1 are shown in Fig. 21-6 after different numbers of simulated injections. The grooves on the machined surface (Fig. 21-6 a) were mostly removed after only 350 simulated injections with zirconia feedstock at 1708C (Fig. 21-6 b). With further increase in the number of simulated injections the grooves were completely removed and after 700 simulated injections (Fig. 21-6 c) the surfaces were smoothed compared with the originally machined surface. The worn surfaces showed fine grooves in the flow direction of the feedstock due to abrasion by the hard zirconia particles and also some synclinal depressions due to washout effects of the flowing feedstock. The number and the depth of these grooves increased with increasing number of simulated injections (Fig. 21-6 c–e). Fig. 21-6 f illustrates the influence of the feedstock temperature. After 1400 simulated injections at 190 8C the grooves from the machining of the specimens prior to testing were still clearly visible. Volumetric wear of the differently heat treated steels C 80 and X38CrMoV5-1 versus Vickers hardness HV30 is shown in Fig. 21-7. There was no significant influence of the hardness detectable for these two steels. However, volumetric

Fig. 21-5 Volumetric wear of the steel X38CrMoV5-1 versus the number of simulated injections (strokes in the test) at an average flow velocity of 245 mm/s with ZrO2 feedstock at temperatures of 170 and 190 8C

21.3 Results

Fig. 21-6 Scanning electron micrographs of the steel X38CrMoV5-1 hardened and tempered at 560 8C: (a) micromachined (micromilled) and after (b) 350, (c) 700, (d) 1400 and (e) 2800 simulated injections with ZrO2 feedstock at 170 8C and (f) after 1400 simulated injections with ZrO2 feedstock at 190 8C (v = 245 mm/s)

wear of the steel C 80 correlated with its microstructure. The least wear was measured on the fine pearlitic microstructure of the normalized steel (Fig. 213 a) and the greatest wear on the specimen tempered at 700 8C showing a coarser microstructure (Fig. 21-3 b). The influence of different feedstocks on the volumetric wear of the steels 30CrMo6, X38CrMoV5-1 and X175VCrMo9-5, electroplated nickel and brass CuZn37 is shown in Fig. 21-8. The least wear was measured for the PM steel X175VCrMo9-5, which showed the highest hardness of all materials (719 ± 9 HV30). There was no significant difference of the volumetric wear measured for the PM steel after the tests with the two different ceramic feedstocks. After tests with zirconia feedstock, the volumetric wear of the other two steels was higher by a factor of about four. There was no distinct correlation between the hardness and the wear resistance of the materials. The high-alloyed steel X38CrMoV5 1 (524 ± 11 HV30) showed significantly more wear than the softer low-alloyed steel 30CrMo6 (454 ± 2 HV30). Volumetric wear of the single-phase reference materials nickel (355 ± 11 HV30) and brass (104 ± 1 HV30) was lower than that of the steels except PM steel X175VCrMo9-5. After 1400 simulated injections with alumina feedstock, the volumetric wear of the electroplated nickel

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21 Tribological Characterization of Mold Inserts and Materials for Microcomponents

Fig. 21-7 Volumetric wear of the differently heat treated steels C 80 and X38CrMoV5-1 versus Vickers hardness HV30 after 1400 simulated injections in the lPIM test with ZrO2 feedstock (T = 170 8C; v = 245 mm/s)

Fig. 21-8 Volumetric wear of the steels 30CrMo6, X38CrMoV5-1 (hardened and tempered at 560 8C), X175VCrMo9-5, electroplated nickel and brass CuZn37 after 1400 simulated injections in the lPIM test with ZrO2 and Al2O3 feedstock (T = 170 8C; v = 245 mm/s)

and the steels except PM steel X175VCrMo9-5 was about double that in the tests carried out with zirconia feedstock. The ranking of these materials with respect to their wear resistance was the same for tests with zirconia and alumina feedstock. The volumetric wear of brass CuZn37 was about six times higher after tests with alumina feedstock than that measured with zirconia feedstock. Worn surfaces of the PM steel X175VCrMo9-5, electroplated nickel and brass CuZn37 are shown in Fig. 21-9 after 1400 simulated injections with zirconia and alumina feedstock. In contrast to the worn surface of the steel

21.3 Results

Fig. 21-9 Scanning electron micrographs of the steel (a, b) X175VCrMo9-5, (c, d) electroplated nickel and (e, f) brass CuZn37 after 1400 simulated injections in the lPIM test with (a, c, e) ZrO2 and (b, d, f) Al2O3 feedstock (T = 170 8C; v = 245 mm/s)

X38CrMoV5-1 (Fig. 21-6), the PM steel X175VCrMo9-5 showed a relatively rough surface with a relief topography due to the selective washout of the softer steel matrix around the hard carbides after tests with zirconia feedstock (Fig. 219 a). The worn surface of the electroplated nickel showed fine grooves and only some minor depressions after the tests with zirconia feedstock (Fig. 21-9 c). The surface of brass CuZn37 was characterized by a scaly topography with small tongue-shaped areas of severely plastic deformation and embedded ceramic particles between them (Fig. 21-9 e). The surfaces of electroplated nickel and brass CuZn37 worn by the alumina feedstock were both characterized by fine grooves in the flow direction of the molten feedstock (Fig. 21-9 d and f). The worn surface of PM steel X175VCrMo9-5 after 1400 simulated injections with alumina feedstock (Fig. 21-9 b) was comparable to that with zirconia feedstock (Fig. 219 a) and characterized by the selective washout of the soft steel matrix around the hard carbide particles.

21.3.2.2 Behavior in Microtribological Sliding Tests Fig. 21-10 shows the friction coefficient as a function of the sliding distance for the self-mated zirconia and alumina sliding pairs at a normal load of 800 mN

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21 Tribological Characterization of Mold Inserts and Materials for Microcomponents

Friction coefficient l

592

Sliding distance s [m]

Sliding distance s [m]

Fig. 21-10 Friction coefficient of self-mated ZrO2 and Al2O3 sliding pairs in (a) air with a relative humidity of 50% and (b) distilled water versus sliding distance (FN = 800 mN, v = 400 mm/s)

and a sliding speed of 400 mm/s in ambient air with a relative humidity of 50% and in distilled water. The friction coefficient of the zirconia sliding pair increased sharply to a maximum value of 0.61 during the first 40 m of the sliding distance in air (Fig. 21-10 a). With prolonged sliding distance, the friction coefficient increased slightly to a quasi-stationary value of 0.65 after about 400 m. The friction coefficient of the unlubricated alumina sliding pair was 0.18 at the beginning of the tests and slowly increased to a quasi-stationary value of about 0.3 after 300 m (Fig. 21-10 a). Compared with the tests in air, both ceramics showed a lower friction coefficient when lubricated with distilled water (Fig. 21-10 b). The water-lubricated zirconia pair was characterized by a short running-in period with a maximum friction coefficient of about 0.5, which decreased with prolonged sliding distance and reached a quasi-stationary value of about 0.35 after 150 m (Fig. 21-10 b). The alumina sliding pair did not show a running-in period under water lubrication but a relatively constant friction coefficient of 0.11 over the total sliding distance (Fig. 21-10 b). The influence of the normal load on the friction coefficient of the self-mated ceramic pairs is illustrated in Fig. 21-11. The unlubricated zirconia sliding pair was characterized by a sharp increase in the friction coefficient at the beginning of the tests and reached maximum values of l between 0.4 (200 mN) and about 0.6 (400 and 800 mN) after about 40 m of sliding distance (Fig. 21-11 a). During the tests at 200 mN, the friction coefficient increased after the running-in period and reached a quasi-stationary value of about 0.59 after 350 m. At a normal load of 400 mN, the friction coefficient of the zirconia pair reached a maximum value of l = 0.72 after about 240 m and decreased with prolonged sliding distance to a value of 0.65 at the end of the tests after 500 m. During the tests with a normal load of 800 mN, the friction coefficient increased from 0.6 after the running in-period to a quasi-stationary value of 0.65 at the end of the tests (Fig. 2111 a). Under water lubrication, the zirconia sliding pair showed a running-in period with maximum values of the friction coefficient in the range 0.5–0.55

21.3 Results

Sliding distance s [m]

Sliding distance s [m]

Sliding distance s [m]

Sliding distance s [m]

Fig. 21-11 Friction coefficient of self-mated (a, c) ZrO2 and (b, d) Al2O3 sliding pairs in (a, b) air with a relative humidity of 50% and (c, d) distilled water at normal loads of FN = 200, 400 and 800 mN versus sliding distance (v = 400 mm/s)

and quasi-stationary values of l = 0.35 (Fig. 21-11 c). There was no significant influence of the normal load on the frictional behavior of the alumina sliding pair both in ambient air and under water lubrication (Fig. 21-11 b and d). During the unlubricated sliding tests, the friction coefficient increased from values between 0.17 and 0.19 at the beginning of the tests to quasi-stationary values between 0.27 and 0.3 after a sliding distance of 500 m at the end of the tests (Fig. 2111 b). The alumina sliding pairs showed under water lubrication no distinct running-in period and a quasi-stationary friction coefficient between 0.11 and 0.13 over the total sliding distance (Fig. 21-11 d). Fig. 21-12 summarizes the influence of the normal load on the quasi-stationary friction coefficient and the amount of linear wear of the self-mated ceramic sliding pairs after a sliding distance of 500 m. A significant influence of the normal load on the quasi-stationary friction coefficient was detectable only with the unlubricated zirconia pair, which showed a slightly increasing friction coefficient with increasing normal load (Fig. 21-12 a). On the other hand, the amount of linear wear measured after a sliding distance of 500 m increased approximately linearly for both zirconia and alumina sliding pairs in unlubricated and

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21 Tribological Characterization of Mold Inserts and Materials for Microcomponents

Normal Force FN [mN]

Normal Force FN [mN]

Fig. 21-12 (a, b) Quasi-stationary friction coefficient and (c, d) amount of linear wear (sum of ball and disc) of self-mated ZrO2 and Al2O3 sliding pairs in (a, c) air with a relative humidity of 50% and (b, d) distilled water after sliding over 500 m at normal loads of FN = 200, 400 and 800 mN (v = 400 mm/s)

water-lubricated contact with increasing normal load. Under all test conditions, wear of the ball specimens was significantly higher than that of the disc specimens, which was typically not measurable. The least ball wear of less than 0.5 lm was measured after unlubricated sliding tests with zirconia at a normal load of 200 mN. However, with increasing normal load the average amount of wear sharply increased to 3.7 lm at 800 mN (Fig. 21-12 c). At a normal load of 200 mN, the unlubricated alumina showed greater wear (Wl = 1.1 lm) than the zirconia but with increasing normal load the linear wear increased less than with the zirconia pair (Fig. 21-12 c). Wear of the water-lubricated zirconia pairs was lower than that of the alumina pairs at all normal loads. After a sliding distance of 500 m the amount of linear wear of zirconia was about 0.6 lm at a normal load of 200 mN and 1.8 lm at 800 mN. Wear of the alumina sliding pairs was between 1.6 lm at 200 mN and 2.7 lm at 800 mN (Fig. 21-12 d). The wear of zirconia was lower (except the tests with 200 mN), but the wear of alumina was greater under water lubrication than in ambient air.

21.3 Results

Fig. 21-13 Scanning electron micrographs of the worn surfaces of (a, b) the ZrO2 disc specimens, (c, d) the ZrO2 balls and (e, f) the Al2O3 balls after 500 m sliding with v = 400 mm/s at a normal load of 800 mN in (a, c, e) ambient air with a relative humidity of 50% and (b, d, f) distilled water

Scanning electron micrographs of worn surfaces of zirconia and alumina specimens are shown in Fig. 21-13 after a sliding distance of 500 m, with a velocity of 400 mm/s and a normal load of 800 mN in air and distilled water. The worn surfaces of the zirconia disc and ball were both covered with layers of densified wear debris after the sliding distance of 500 m in air (Fig. 21-13 a and c). Locally, flaking of these layers occurred. The grooves on the machined surface of the zirconia disc specimen were clearly visible after the sliding distance of 500 m in distilled water and wear particles in the form of small rolls on the disc surface indicated the detachment of thin tribochemical films (Fig. 21-13 b). The worn surface of the zirconia ball was smooth after the tests in distilled water compared with that run in air and only a few loose wear particles were found (Fig. 21-13 d). The worn surfaces of the alumina balls showed a relief topography due to the different wear rates of individual alumina grains depending on their crystallographic orientation after tests in both air and distilled water (Fig. 21-13 e and f).

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21.4 Discussion 21.4.1 Wear of Mold Inserts in Micropowder Injection Molding Test

The laboratory tests identified that in microinjection molding using ceramic feedstocks, the high injection pressure and flow velocity of the molten feedstock typically result in abrasive and/or erosive wear of the mold inserts. In order to verify these experimental results, different mold inserts were analyzed after practice in lPIM with ceramic and metallic feedstocks at the Institute for Materials Research III at the Forschungszentrum Karlsruhe (FZK-IMF III). Fig. 2114 shows scanning electron micrographs of mold inserts made from high-alloyed steel X38CrMoV5-1, electroplated nickel and brass CuZn39Pb3. The worn surfaces of the mold inserts were grooved by the hard particles from the molten feedstock and the machined surfaces were partially smoothed (Fig. 21-14 b, d and f). Especially the surface of the mold insert made from electroplated nickel

Fig. 21-14 Scanning electron micrographs of mold inserts made from (a, b) the steel X38CrMoV5-1, (c, d) the electroplated nickel and (e, f) the brass CuZn39Pb3 used for micropowder injection molding with ceramic and metallic feedstock (FZK-IMF III)

21.4 Discussion

also showed a wavy surface topography due to local washout of material (Fig. 21-14 d). The worn surfaces after the laboratory tests were characterized by synclinal depressions, a wavy surface topography and fine grooves depending on the composition of the feedstock and the material of the mold insert specimens. The results from this research work are in good agreement with results from previous experiments which showed that hardness is not a sufficient criterion under the conditions of erosive wear to evaluate and compare the wear resistances of different materials [21, 24, 25, 50–55]. Important influences in addition to hardness are microstructural parameters of the materials such as homogeneity and bond strength between different phases and the characteristics of the feedstock (e.g. amount, size, shape and hardness of the hard particles). Additional short-time experiments were performed using specimens with polished surfaces to obtain a better understanding of the different wear behaviors observed for the multiphase steels, the single-phase electroplated nickel and the brass during the laboratory tests. Fig. 21-15 shows worn surfaces of the tool steel, electroplated nickel and brass CuZn37 after one and 11 simulated injections with zirconia feedstock. After only one simulated injection many short, comma-shaped grooves were visible on the surface of the tool steel X38CrMoV5-1 (Fig. 21-15 a). These small grooves were extended to small depressions with increasing number of simulated injections (Fig. 21-15 b). The inhomogeneous material removal was attributed to the microstructure of the steel, which was characterized by a small amount of hard carbides (< 5 vol.%) embedded in a soft steel matrix. It has been shown previously that hard phases within a steel can increase wear resistance only if the length of the free path between two adjacent hard particles is not clearly larger than the particle size of the attacking abrasive particles [24]. Small amounts of inhomogeneously distributed hard carbides in a softer matrix can even reduce wear resistance although they might increase strength and hardness. This was reported in [51] from experiments with upper bainitic austempered ductile iron. The positive effect of a homogeneous and fine-grained microstructure was also illustrated by tests with the differently heat treated steels C 80 and X38CrMoV5-1 (Fig. 21-7). The amount of volumetric wear was not correlated with the hardness but with the microstructure of the steel C 80. The least wear was measured for the fine pearlitic microstructure of the normalized steel (Fig. 21-3 a) and the greatest wear for the steel C 80, which was tempered at 700 8C, resulting in a coarser microstructure (Fig. 21-3 b). Comparable results on the positive influence of a fine microstructure on wear resistance under erosive wear conditions have been reported for some low-alloyed steels [53] and chrome white irons [54]. The significantly lower volumetric wear of the electroplated nickel compared with the steel C 80 of about the same hardness also exemplifies the positive influence of a homogeneous microstructure. The Vickers microhardness for the more wear-resistant nickel (380 ± 15 HV50) showed a much smaller variation than that of the steel C 80 (416 ± 43 HV50). This resulted in homogeneous grooving and material removal during the laboratory tribotests with electroplated nickel (Fig. 21-15 c and d).

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Fig. 21-15 Scanning electron micrographs of (a, b) the steel X38CrMoV5-1 (hardened and tempered at 560 8C), (c, d) the electroplated nickel and (e, f) the brass CuZn37 after (a, c, e) one and (b, d, f) 11 simulated injections in the PIM test with ZrO2 feedstock (T = 170 8C, v = 245 mm/s)

The wear behavior of the brass CuZn37 was completely different from that of all other materials in tests with zirconia feedstock. The roughness of the surface was significantly increased after only one simulated injection with zirconia feedstock (Fig. 21-15 e). The wavy surface topography was already established after 10 more simulated injections (Fig. 21-15 f) and developed more and more with increasing number of simulated injections (Fig. 21-9 e). It was assumed that the rough wavy surface topography resulted from swirls of the molten feedstock which led to locally different stress conditions [50]. The brass could be protected from severe wear due to the embedding of zirconia particles into the surface, but the rough surfaces should result in problems during demolding of solidified microcomponents. The results from tests with alumina feedstock demonstrated that the wear behavior was also distinctly influenced by the characteristics of the feedstock. Tests with alumina feedstock resulted in dramatically increased wear compared with the tests with zirconia feedstock for all materials except the PM steel. This could be attributed to the greater hardness of the alumina (ca. 2000 HV) compared with the zirconia (ca. 1200 HV) and to the sharp-edged particle shape. This resulted in greater abrasive wear and therefore in smaller differences in wear re-

21.4 Discussion

sistance of the materials. Furthermore, the higher volume fraction (58 vol.%) of alumina particles within the feedstock resulted in a higher viscosity of the feedstock and therefore increased abrasiveness [10]. On the other hand, a slightly increased temperature of the molten feedstock (190 instead of 170 8C) led to a significantly lower viscosity of the feedstock and therefore to a reduced amount of volumetric wear. 21.4.2 Friction and Wear During Microtribological Sliding Tests

Friction and wear of zirconia and alumina specimens produced by lPIM were characterized in unidirectional sliding contact with regard to their utilization as tribologically high-loaded microcomponents. Fig. 21-16 shows the volumetric wear coefficient of the self-mated pairs calculated from the amount of linear wear measured after a sliding distance of 500 m with v = 400 mm/s at a normal load of 800 mN as a function of the friction coefficient. Alumina showed a lower friction coefficient than zirconia in both unlubricated and water-lubricated sliding contact. This was attributed to the formation of a soft aluminum hydroxide film in the contact zone due to tribochemical reactions. The low shear strength of the aluminum hydroxide can effectively reduce friction and thereby can lead to a low wear rate of self-mated alumina in sliding contact [38–42]. The lower friction coefficient but higher wear rate of alumina under water lubrication could be explained by the enhanced rate of formation and dissolution of the hydroxide in an aqueous solution compared with humid air [39, 41]. Zirconia showed the highest wear coefficient in unlubricated sliding contact and a significant dependence of the wear rate on the normal load (Fig. 21-16; see also Fig. 21-12). There was no indication of the formation of a soft tribochemical film comparable to that on alumina. This resulted in high friction and a large amount of small wear particles (Fig. 21-13 c). Under water lubrication, self-mated zirconia was characterized by a distinct running-in period with a significant decrease in the friction coefficient. Similarly to alumina, tribochemical hydroxide films can form in the contact zone of self-mated zirconia tribo cou-

Fig. 21-16 Wear coefficient kV versus friction coefficient of the self-mated ZrO2 and Al2O3 sliding pairs (FN = 800 mN, v = 400 mm/s, s = 500 m)

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ples [37, 38]. Wear particles in the form of small rolls (Fig. 21-13 b and d) are characteristic of the delamination of such a thin tribochemical film. As the rate of formation and dissolution of the hydroxide is typically lower for zirconia, the wear rate was lower but the friction coefficient was higher compared with those of alumina under water lubrication. Under the operating conditions chosen for the tests reported here, there was no evidence for the hydrothermal destabilization of the metastable tetragonal phase by water, which was described elsewhere [46–48].

21.5 Conclusions

The wear behavior of the unalloyed, low- and high-alloyed steels C 80, 30CrMo6, X38CrMoV5-1 and X175VCrMo9-5, electroplated nickel and brass Cu63Zn37 used as mold inserts in lPIM was studied in a laboratory tribometer simulating powder injection molding with zirconia and alumina feedstock at 170 and 190 8C. The results indicated that in addition to hardness microstructural parameters such as homogeneity in single-phase materials (nickel and brass), homogeneous distribution of hard particles in multiphase materials (steel) or the ability of work hardening and plastic deformation influence the wear resistance of the tested materials to a high degree. The wear behavior was also strongly influenced by the characteristics of the feedstock and the test parameters (e.g. temperature). The highest wear resistance was obtained with the high-alloyed PM steel X175VCrMo9-5, which showed a high hardness due to about 20 vol.% of fine, homogeneously distributed carbides within a martensitic matrix. Microtribological tests under unidirectional sliding conditions were carried out on self-mated TZP-ZrO2 and Al2O3 pairs in air with a relative humidity of 50% and under water lubrication. The results showed that both zirconia and alumina can be favorable materials for tribologically high-loaded micromechanical components, particularly if lubricated with water. Alumina showed lower friction but higher wear under water lubrication compared with the unlubricated tests, whereas zirconia showed both reduced friction and wear under water lubrication. Further investigations to characterize the qualification of the ceramics under slip-rolling conditions are in progress.

21.6 Acknowledgments

The authors thank the Deutsche Forschungsgemeinschaft (DFG) for financial support within the frame of the Center of Excellence in Research 499 ‘Construction, production and quality assurance of molded microcomponents made from metallic and ceramic materials’ and Mr. Helmut Iwanek for his assistance with the experimental work.

21.7 References

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53 A. Sundström, J. Rendon, M. Olsson,

Microsystem Technology; FZKA 6620, Karlsruhe: Forschungszentrum Karlsruhe, 2001. 50 H. Uetz, Abrasion and Erosion; Munich: Carl Hanser, 1986. 51 F. Y. Hung, L. H. Chen, T. S. Lui, ‘A study on the particle erosion of upper bainitic austempered ductile iron,’ Wear 2002, 252, 985–991. 52 U. Beste, L. Hammerström, H. Engqvist, S. Rimlinger, S. Jacobsen, ‘Particle erosion of cemented carbides with low Co content,’ Wear 2001, 250, 809–817.

‘Wear behaviour of some low-alloyed steels under combined impact/abrasion contact conditions,’ Wear 2001, 250, 744–754. 54 R. J. Llewellyn, S. K. Yick, K. F. Dolman, ‘Scouring erosion resistance of metallic materials used in slurry pump service,’ Wear 2004, 256, 592–599. 55 I. Hussainova, ‘Effect of microstructure on the erosive wear of titanium carbidebased cermets,’ Wear 2003, 255, 121– 128.

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22 Development of a Simulation Tool for Wear in Microsystems V. Hegadekatte, Institute for Reliability of Components and Systems, University of Karlsruhe (TH), Germany N. Huber, Institute for Material Research II (IMF II), Forschungszentrum Karlsruhe, Germany O. Kraft, Institute for Reliability of Components and Systems, University of Karlsruhe (TH), and Institute for Material Research II (IMF II), Forschungszentrum Karlsruhe, Germany Abstract Wear is one of the major reliability issues of micromachines owing to their inherent high operating frequencies and their high surface-to-volume ratio. Experimental methods, such as pin-on-disc, scratch test or atomic force microscopy, are not sufficient for the quantitative prediction of the progress of wear and the life span of micromachines. The strategy discussed in this chapter makes use of Archard’s wear model, which is implemented within a finite element-based wear simulation tool. This approach allows for simulating pin-ondisc wear experiments with the goals of achieving a better understanding of the phenomena of the wear process and of determining the material parameters relevant to wear. It will be demonstrated that our approach has the potential to predict wear for small-scale complex tribo-systems with high operation frequencies.

Keywords wear; wear modeling; wear simulation; micromachines; microsystems

22.1 22.2 22.2.1 22.2.2 22.3 22.4 22.5 22.6 22.7

Introduction 606 Modeling and Simulation of Wear in the Literature FEM Implementation 609 Post-Processor 609 Development of the Wear Simulation Tool 610 Wear Simulation Using the Wear-Processor 613 Conclusion 622 Acknowledgment 622 References 623

Advanced Micro and Nanosystems Vol. 4. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31493-8

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22 Development of a Simulation Tool for Wear in Microsystems

22.1 Introduction

Micromachines promise to revolutionize our present environment by facilitating the development of a wide variety of smart products. One of the main advantages of micromachines is their quick response time owing to their extremely low inertia. However, their very high surface-to-volume ratio adversely affects their reliability as they often fail due to excessive wear. Wear being a surface phenomenon is identified as a critical factor that can limit substantially the life span of micromachines. It is compounded by the fact that the surface forces become significant compared with the volume forces at these extremely small dimensions. Investigations on microturbines made of polycrystalline silicon showed that the typical life span of such components was around 1 ´ 106 cycles at 5000 revolutions per second [1]. In [2], a comparison of a well-designed watch with a MEMS (MicroElectro Mechanical System) showed that 1 min of usage of a MEMS represents an amount of wear corresponding to a degradation that occurs in more than 10 years in a well-designed watch bearing. These examples show that the tribological behavior plays a key role in the performance of micromachines. A promising approach to address the improvement of the wear resistance in MEMS is to employ ceramic materials as described in Chapter 21, but predicting wear is still essential for the development of reliable devices. At present, in situ wear measurements are the most realistic methods to predict wear in microcomponents. However, the manufacture of prototypes is highly expensive in terms of both time and money. Often experimental techniques such as the pin-on-disc test, scratch test and atomic force microscopy (AFM) are used to characterize the tribological properties of various materials used in microsystems technology. They are not sufficient, however, for an exact prediction of the progress of wear in micromachines and their useful life span. As an example, in part B of [3], the design and production of a micropump were discussed including the prediction of failure [4] and wear ([3], pp. 83–87). One important outcome of this work is that there exists no adequate predictive method for wear in micromachines with continuously changing contact. Therefore, it is appropriate to develop a simulation tool, a Wear-Processor, which will be described in detail in Section 22.3. This approach is very lucidly represented in Fig. 22-1. Our strategy is twofold, namely: (i) identify the wear model by comparing the results from the simulation of wear with the experimental results of, e.g., a pin-on-disc problem (see Chapter 21) and (ii) use this identified wear model to predict wear and wear parameters in real systems with the goal of estimating the life span of a micromachine, such as the SFB 499’s demonstrator (see Chapters 1 and 2). The approach applies to both the macro- and especially the microscales. However, it is of special interest to microscale applications that the high contact pressures and sliding velocities encountered in micromachines would result in high wear rates in the relevant microtribological experiments. There exists hardly any possibility to improve the finishing of complex-shaped microparts, e.g. microgears, for adjusting the tolerances and the surface quality,

22.2 Modeling and Simulation of Wear in the Literature

Fig. 22-1 Strategy for wear prediction in micromachines

which is another reason for the importance of a simulation tool that could predict the influence of the production process and its achievable tolerances on the lifetime of a micromachine. Using the Wear-Processor to identify a wear model and subsequently the included parameters by simulating microtribology experiments conducted within the parameter space of a certain micromachine, would restrict such a wear model for simulating wear on this particular system. Additionally, special care has to be taken to mesh the microcomponents, taking into account the high aspect ratios. With the wear simulation tool to be discussed in this chapter, it is attempted to close the gap between in situ wear measurements, standard tribological experiments (Chapter 21) and the actual operation of the micromachine (Chapter 2). A satisfactorily working wear simulation tool would be able to provide guidelines for material selection (Section 3) and microsystem design (Section 1). Such a tool will also be able to spell out the requirements on the materials and production technology (Section 3), so that a given life span can be achieved. The simulation tool will be able to determine the loss of material at the surface and in this way the continuous change of the kinematics, e.g. to consider the microplanetary gear train (SFB 499’s demonstrator) for a more realistic stress analysis. Hence the life span can be predicted more accurately both by the failure due to kinematics and the breakdown (see Chapters 19 and 20) due to drastically increased loads resulting from wear. In addition, noise and vibration characteristics of micromachines can be studied more realistically and accurately.

22.2 Modeling and Simulation of Wear in the Literature

It is well known that the phenomena of wear are among the most complex and difficult to model of all mechanical events, primarily owing to the complex structure of engineering surfaces, severe deformation in the contact region, frictional heat generation, presence of contaminants and wear debris, lubrication,

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atomic range interactions, chemical reactions at the contact surfaces, etc. (see [5] for a detailed treatment of sliding wear). Owing to the challenges that wear offers and its extreme importance to engineering, it has been a subject of intense scientific and empirical investigations with mechanistic insights that can qualitatively explain a wide variety of wear phenomena reasonably well [6]. The literature is rich with wear equations, e.g. see [7, 8] where nearly 300 wear/friction models are identified involving an enormous range of material and operating parameters. In [9], a very elaborate treatment of sliding friction can be found. In spite of the large number of wear models in the literature, no model can predict wear with reasonable accuracy a priori based only on material properties and contact information [10]. Owing to design restrictions in micromachines, sliding wear is the most commonly encountered type of wear. It is mechanistically more complex than certain other forms of wear because it involves not only the cutting and plowing included in abrasive wear but also the adhesion of asperities, third bodies, subsurface crack initiation and growth, the transfer of material to and from the mating surfaces, subtle changes in surface roughness during running in, tribochemical film formation and many other processes [11]. Owing to the above complexities, it would be a formidable if not impossible task to derive a wear equation from first principles that describes sliding wear satisfactorily and that is truly predictive in nature. Archard was the first to propose a phenomenological wear model to describe sliding wear [12]. It assumes that the critical parameters in sliding wear are the stress field in the contact and the relative sliding distance between the contacting surfaces. The constant of proportionality in Archard’s wear model, commonly termed Archard’s wear coefficient, has been interpreted in various ways in the literature, depending on the preferences of the various authors (e.g. see [13]), namely: (i) as the fraction of the asperities yielding wear particles, (ii) as the ratio of the volume worn to the volume deformed, (iii) as a factor inversely proportional to a critical number of load cycles, (iv) as a factor reflecting the inefficiencies associated with the various processes involved in generating wear particles, (v) as the number of repeated asperity encounters to produce rupture, etc. Hardness is the only material property appearing explicitly in the linear wear equation given by Archard, whereas the effects of the other material properties are included in the wear coefficient. It is this simplicity of Archard’s wear model that makes it attractive as a starting model in more complex analysis for the development and improvement of wear models which will be discussed next. Mild wear can be described fairly accurately and conveniently on a global scale with Archard’s wear model in spite of the fact that it is phenomenological in nature and gives little insight into the dominant wear mechanisms. Most experimental investigations try to study the wear coefficient as a measure of the material’s performance against wear under the given conditions on a global scale. The coefficient of wear is often assumed to be constant, but it is often observed to change by several orders of magnitude when transitions occur in response to changes in experimental variables such as load, speed and environment [14]. The role of hardness is reasonably well de-

22.2 Modeling and Simulation of Wear in the Literature

fined for abrasion but is much less clear in sliding situations. Several attempts to improve Archard’s wear model have been made: (i) the delamination theory of wear is based on the behavior of dislocations at the surface, subsurface crack and void formation and subsequent joining of cracks by shear deformation of the surface [15]; (ii) modifications of Archard’s wear equation for computing wear in highly elastic/pseudo-elastic materials are given in [16]; and (iii) wear equations for polymers have been developed based on a dimensional analysis [17], while polynomials have been fitted to experimental curves [18, 19], which makes them valid only within the experimental range. More recent developments have concentrated on developing computer-based simulation techniques, which are made possible by the advent of modern highperformance computers and the rapid development of methods for solving nonlinear problems with finite element (FE) analysis. The research on simulating wear from the available literature can be divided broadly into two approaches, namely the FEM implementation and the use of a post-processor for the computation of wear using a suitable wear model with or without the use of FEM, which will be discussed in the next two sections. 22.2.1 FEM Implementation

The first approach is to implement directly a wear model in the material law of an FE analysis. In [20], a modification of Archard’s wear model is implemented where the hardness of the softer material is allowed to be a function of temperature; other features such as surface evolution due to wear, finite deformation, thermoplasticity and frictional contact were included. In [21], a user-defined contact and wear element is implemented that activates the appropriate predominant wear mechanism depending on the sliding velocity and the normal load from the wear mechanism maps of steel on steel as developed in [22], which makes it specific to the testing conditions of the wear maps. A numerical model for simulating the frictional wear behavior is presented within a fully non-linear kinematic setting, including large slip and finite deformations [23]. The main drawback of this approach is the considerable computational expense. 22.2.2 Post-Processor

A faster and efficient approach is post-processing of the FE results obtained from the solution of a general contact problem with a suitable wear model to compute the progress of wear for a given time interval/sliding distance. In [24– 26], a computational approach to predict sliding wear is proposed based on the theory that owing to repeated asperity contact, there is accumulation of plastic strain with every loading event (ratcheting). When the accumulated strain exceeds a critical value, the material is said to have failed or worn out. This theory was developed in [27]. Even though the work reported in [24–26] provides novel

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approaches to simulate wear, the assumptions and simplifications are considerable. Some of the approximations are as follows: (i) assumption of plane strain conditions and therefore solving the contact problem in 2D; (ii) the contact is between a rigid and a deformable surface and so the solution is obtained only for the deformable surface; and (iii) lack of consideration or idealization of the rapidly changing contact conditions as a result of wear. The wear simulations in [28–31] are based on Archard’s wear model and are implemented as an FE postprocessor on 2D problems. Similarly, the studies in [32, 33] are based on the Archard’s wear model and make use of the elastic foundation method [34] for the computation of the contact pressure instead of FE analysis in order to reduce the computational time. So far we have given a brief survey of the research on modeling and simulation of wear from the literature. It can be noted from this section that computational approaches have been the method of choice for the more recent advances in simulating wear. However, for simulating wear in microsystems, Archard’s wear law is the most popular model as in [2], where it is used to predict wear in rotating pivots for MEMS. A microcontact and wear model is developed for predicting the material removal rate from silicon wafer surfaces during chemical–mechanical polishing in [35], the developed equation being a representation of Archard’s wear law. In [36], a simulation scheme based on Archard’s wear model is developed for the surface shape and contact pressure evolution during copper chemical–mechanical polishing. At a more fundamental level, the atomic processes occurring at the interface of two materials when brought in contact, separated or moved with respect to one another was studied with scanning probe microscopy and molecular dynamics simulations in [37]. Such methods will greatly aid in the fundamental understanding of the interactions at an interface. In the following part of this chapter, we will discuss the development of the Wear-Processor, which implements Archard’s linear wear model to simulate dry sliding wear on a pin-on-disc problem. This tool works in association with a commercial FE package for the computation of the stress distribution. As we have seen so far in this chapter, the phenomena of wear are extremely complex and, in order to simulate wear in micromachines confidently, we need to improve the existing wear model and subsequently determine the included material parameters, which forms the future course of work with the Wear-Processor.

22.3 Development of the Wear Simulation Tool

In this section, the functioning of the wear simulation tool ‘Wear-Processor’ is presented in a concise way. The Wear-Processor employs FEM for the computation of the contact stresses using the commercial package ABAQUS [38] because of its user friendliness for solving contact problems. However, in order to predict wear and the life span of complex systems, several hundred thousand

22.3 Development of the Wear Simulation Tool

operating cycles have to be simulated. Under these circumstances, an FE postprocessor is the most logical choice considering the magnitude of the computational expense of such large-scale contact simulations. Also, the Wear-Processor does not aim to simulate the entire sliding process of a pin-on-disc experiment with dynamic FE analysis but instead treats the problem of sliding wear as ‘quasi-static’ in order to save on the computational expense. Therefore, to be as realistic as possible, a general contact problem involving two deformable bodies is solved a number of times at different stages (wear steps) of the sliding process with the surface evolving through different wear steps. In Fig. 22-2, the entire process inside the dashed line represents the Wear-Processor. The process begins with the FE solution of the contact problem, which involves the explicit modeling of the contact geometry, selecting a material model, supplying the coefficient of friction and applying the necessary boundary condition. The working of the Wear-Processor begins with the reading of the binary coded results from the contact simulation with an interface program (FE-Post). The inward surface normal vectors at each of the surface nodes are computed based on the element topology for the 3D geometry. The normal direction is computed by taking the cross product of the four edge vectors that are connected to the node concerned [39]. The contact pressure at each of the surface nodes is computed from the normal vector and the stress tensor averaged at each surface node using tj ˆ rij ni

…1†

where tj is the traction vector, rij is the stress tensor, ni is the inward surface normal vector at the corresponding surface node, the subscripts i and j correspond to the tensor components in 3D space depending on the contact situation and then the contact pressure p at each surface node is calculated using p ˆ tj nj

…2†

The Wear-Processor implements Archard’s wear model to calculate linear wear at each of the surface nodes. However, such an implementation does not explicitly take the effect of the wear debris as all these effects will be implicitly included in the wear coefficient. The wear model is implemented on the local scale, i.e. it considers the current contact pressure and the location of each node and is given by VW =s ˆ kD FN

…3†

where VW is the wear volume, s is the sliding distance, kD is the dimensional wear coefficient and FN is the applied normal load. The linear wear h is calculated by dividing Equation (3) on both sides by the real area of contact: h=s ˆ kD p

…4†

611

Fig. 22-2 Flow chart of the Wear-Processor: utotal is the total displacement, which is the algebraic sum of the elastic displacement (uelastic) and the displacement computed due to wear (uwear), s is the current sliding distance and smax is the maximum sliding distance. PATRAN is a commercial FE pre- and post-processor

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22.4 Wear Simulation Using the Wear-Processor

where p is the contact pressure at each surface node. In the Wear-Processor, the above wear law is discretized with respect to the sliding distance as hj ˆ kD pDsj ‡ hj

1

…5†

where hj is the total linear wear up to the jth wear step, hj–1 is the total linear wear up to the (j–1)th wear step and Dsj is the increment of sliding distance for the jth (current) wear step. A specific case is the computation of wear for a pin-on-disc experiment, as shown in Fig. 22-3. The pin surface nodes are in contact with the disc through the entire sliding process and, therefore, the computation of wear is straightforward. However, for the disc surface nodes, which come in contact with the pin only once per revolution, the wear has to be computed only for this contact duration. The contact pressure at the disc surface nodes of the FE mesh has to be integrated over the entire wear track at each of the disc wear track radii. The increment of the linear wear on the wear track can be calculated from the integration of the contact pressure over the entire wear track. An efficient way to accomplish this computation is presented in [39]. Each surface node is then shifted in the direction of the inward surface normal depending on the amount of wear at the corresponding surface node in the current wear step. In order to maintain the regularity of the FE mesh and also not to be limited by the surface element height, a mesh correction method (‘Re-meshing’ in Fig. 22-2) is implemented. It is accomplished by a boundary displacement method given in [40]. Once the linear wear at a surface node in a particular wear step has reached a certain percentage (denoted by d in Fig. 22-2) of the corresponding surface element height in the current wear step, the wear processing is stopped in order to re-mesh the model. The nodal coordinates, added to the respective displacements after the re-meshing, will form the reference configuration for the next wear step. The size of the wear step (the amount of sliding performed in each wear step) directly affects the accuracy of the wear processing. The larger the size of the wear step, the larger will be the error, owing to the considerable variation in the contact condition (contact area and the contact pressure field). On the other hand, the finer the size of the wear step, the higher will be the number of contact simulations required, which forms the bottleneck as far as the computation time is concerned. To visualize the progress of wear, the displacement due to wear and the elastic displacements are combined for each wear step and written as an ABAQUS compatible file (Wear-Post) that can be read with a general FE post-processor.

22.4 Wear Simulation Using the Wear-Processor

In principle, because of the generality of this method, wear on components for any material combination can be simulated. For a first demonstration the WearProcessor is applied for simulating wear on a loaded spherical steel pin sliding

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over a steel disc (‘pin-on-disc’) in this section. The material properties of the pin and the disc used in the simulation are as follows: Young’s Modulus, E = 200 GPa and Poisson’s ratio, m = 0.29. The geometric details and the discretization of the pin in the contact region (also a similar discretization for the disc surface contact region is used) of the FE model is shown in Fig. 22-4. Here, Rp represents the radius of curvature of the pin. In Fig. 22-4, the FE model is built as a very small slice of the pin-on-disc from that part of the entire pin-on-disc model which is within the dashed circle in Fig. 22-3. For the wear computation the following values for the notations in Fig. 22-3 were used: FN = 20 N, RP = 1.5 mm, RWT = 3 mm, RD = 10 mm and tD = 2 mm. The top surface of the pin is loaded with a force of 20 N in the form of a distributed load (pressure). For the FE analysis, the coefficient of friction l = 0.5 is assumed. For the FE contact analysis, the loaded pin is slid by an element length in the z-direction (see Fig. 22-3) which is infinitesimal compared with the radius of the wear track, in order to include the effects of friction. An elastic material law and a deformable-deformable contact were used in the FE simulations. Pin-on-disc is an interesting (and experimentally important as shown in Chapter 21) 3D sliding contact problem because the initial contact is Hertzian and it begins to conform (non-Hertzian contact) with increasing sliding distance as a result of wear. The value for the dimensional wear coefficient, kD = 2 ´ 10–13 mm3/N mm, is the same for both the steel pin and the steel disc, and is several orders of magnitude lower than the values found in Chapter 21. The reason for the choice of this very low wear coefficient was to simulate a few thousand revolutions of the pin on the disc in each wear step, thus simplifying the handling of the wear data during the development of the Wear-Processor. However, a very low value of the wear coefficient is not totally unrealistic, as in [41] a wear coefficient of roughly kD = 9.6 ´ 10–13 mm3/N mm was used in a failure model for poly-Si micromachines. The evolution of the contact pressure and the corresponding

Fig. 22-3 Model of a spherical loaded pin revolving over a disc in dry sliding contact. FN is the applied normal load, RP is the radius of the pin, RWT is the radius of the wear track, RD is the radius of the disc and tD is the thickness of the disc. The geometry inside the dashed line is used for the FE simulation by the Wear-Processor (see Fig. 22-4 for the FE model)

22.4 Wear Simulation Using the Wear-Processor

Fig. 22-4 FE Model of the pin-on-disc (top) and the discretization of the mesh in the contact zone (enlarged) is shown by the 2D template for the 3D mesh for the pin (bottom)

evolution of the stresses in the direction of sliding as the sliding progresses can be seen in Fig. 22-5 a–f. The surface plots are shown at the end of three intermediate wear steps in and around the contact region on the disc surface. It can be seen from Fig. 22-5 a, c and e that the contact pressure decreases progressively with the sliding distance, owing to the noticeable increase in the contact area. Fig. 22-6 a shows a graph of the stress in the normal direction to the contact plotted against the location across the sliding direction. The stress distributions plotted for the initial and final configuration (after 1.93 ´ 105 revolutions of the loaded spherical pin over the disc) correspond to the surface plots in Fig. 22-5a and e. For reference, the contact pressure values obtained from the ‘Hertz Formula’ described elaborately in [42] are also plotted in Fig. 22-6 a. The results in Fig. 22-6 b are essentially the same, as shown in Fig. 22-5 b and f in the form of a graph for the initial and final configurations. It should be noted that the graphs are plotted at the surface nodes on the disc surface in the contact region. The values obtained from the analytical solution (see [42] for further details) are also plotted on the same graph. There is good agreement between the FE simulated and the analytical solution, proving the suitability of the applied boundary conditions for the FE analysis. The analytical computation of the stresses is shown only for the initial configuration (before wear) for the spherical pin and the flat disc with elastic material properties. In Fig. 22-6 c, the displacement, uyy, in the normal direction to the contact before and after wear is plotted. It can be seen from Fig. 22-6 d that as the sliding progresses, the increase of the contact area leads to a decreasing pressure in the contact region. Fig. 22-6e shows the surface plot of the linear wear, h, on the disc surface in and around the contact region. It can be seen from Figs. 22.3 and 22.4 that the FE model of the disc does not comprise the entire disc. Instead, only a small re-

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a

b

c

d

e

f Fig. 22-5 Distribution of normal stress ryy and tangential stress rzz plotted on the disc surface nodes in the contact region at different stages of sliding (wear steps). (a) ryy after the initial contact (before wear); (b) rzz after the initial contact (before wear); (c) distribution of ryy after 7.9 ´ 104 revolutions; (d) distribution of rzz after 7.9 ´ 104 revolutions; (e) distribution of ryy after 1.93 ´ 105 revolutions; (f) distribution of rzz after 1.93 ´ 105 revolutions. The direction of sliding is along the z-axis

22.4 Wear Simulation Using the Wear-Processor

a

b

c Fig. 22-6 (a) Graph of normal stress, ryy, vs. x-coordinate at the surface nodes in the contact region of the disc; (b) graph of tangential stress rzz vs. z-coordinate at the surface nodes in the contact region of the disc; (c) graph of normal displacement, uyy, vs. x-coordinate plotted at the surface nodes in the contact region of the disc

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d

e

f

Fig. 22-6 (d) graph of contact area vs. number of revolutions (sliding distance) of the pin over the disc; (e) surface plot of linear wear, h, on the disc surface near the contact region; (f) graph of progress of wear vs. number of revolutions (sliding distance) of the pin over the disc

22.4 Wear Simulation Using the Wear-Processor

presentative portion of the disc is considered in order to reduce the computation time. Hence the wear computed for each node on the FE mesh of the disc surface is representative of any location with the same radii on the entire disc. However, one can appreciate the glimpses of the curvature in the wear track in the surface plot shown in Fig. 22-6 e. The progress of wear over the sliding distance (number of revolutions) on the pin and the disc surfaces is shown in Fig. 22-6 f. It can be seen from Fig. 22-6 f that the slope of the wear curve for the pin is steadily decreasing, because the severity of wear decreases as the contact conforms and so the linear wear per unit sliding distance continuously decreases (running-in). The difference in the amount of wear on the pin and the disc surface is simply due to the fact that each point on the pin is in contact for a much longer time than the points on the disc surface in the wear track. The numerical roughening of the evolved surface, as seen in Figs. 22.5 a–f and 22.6 a and b, is because of numerical restrictions related to the discretization. A finer FE mesh and a smaller wear step size can help in reducing the problem. However, such a strategy increases the computation time enormously. These results can be qualitatively compared work the experimental work in Chapter 21 and in [30] and [43–46]. In [43], it was found that the wear on the disc was less than onetenth of that of the pin and in the other studies cited the wear on the disc was negligible. In [46], the cross-section of the evolved wear track can be qualitatively compared with Fig. 22-6 e. In Fig. 22-7, the progress of linear wear on the pin surface obtained from the various wear models is plotted over the number of revolutions (sliding distance).

Fig. 22-7 Comparison between the wear on the pin surface from various wear models: the Wear-Processor (local wear model), global wear model considering the initial maximum pressure (GWM – In. Max. Pre.), global incremental wear model considering the maximum pressure (GIWM – Max. Pre.) and global incremental wear model considering the average pressure (GIWM – Avg. Pre.)

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In experimental investigations of wear, the contact pressure computed from the initial configuration (Hertzian peak pressure) is usually used to identify Archard’s wear coefficient, i.e. a global wear model is identified as shown in Fig. 22-7. When using the Wear-Processor, the current contact pressure, location and surface topology are considered in the computation. We can see that the global wear model considering the peak pressure is a linear curve with a slope represented by the product of the dimensional wear coefficient and the initial peak pressure. The curve for the linear wear obtained from the Wear-Processor has the same initial slope and then progresses non-linearly (running-in), as is qualitatively observed in experiments [29, 30]. Usually the wear coefficient is computed at the end of the experiment by fitting the entire sliding range. The geometry, which is specific to a particular tribo-system, is then somehow included in the wear coefficient, and also the decay of the contact pressure as the wear progresses is included implicitly in the fit. However, if we are able to identify a wear model using the Wear-Processor fitting the experimental curve, then the derived wear coefficient should be dependent only on the material and tribological parameters. A wear model identified in this way can be more accurately applied to any general tribo-system, since the actual geometry is not contained in the wear coefficient, but is considered by the FE simulation. Once the wear model is identified, it can be used more reliably to predict wear on a different geometry, provided that the pressure and the velocity are in the same range as in the experiments from which the wear model was identified. One of the drawbacks of the global wear model discussed above is that the continuous decrease in the contact pressure is not considered explicitly. However, this drawback can be handled by considering a global incremental wear model. The term ‘incremental’ means that the contact pressure is updated at different intervals of sliding. The assumptions in this model are as follows: (i) wear takes place only on the pin; (ii) the worn out surface is flat; (iii) either an average pressure (load divided by area) or a maximum pressure on the contacting surfaces is considered for the computation of wear; and (iv) frictional effects are not considered. The computation using the global incremental wear model considering the average pressure is accomplished as shown in Fig. 22-8. The elastic displacement, he, is computed according to [47]. For the curve representing the global incremental wear model considering the maximum pressure in Fig. 22-7, the maximum pressure is computed as 1.5 times the average pressure as is the case for the initial Hertzian contact. The non-linear curve from the global incremental wear model considering the maximum pressure also has the same initial slope, because all the three curves discussed so far initially start with the Hertzian peak pressure for the computation of wear. The curve for the global incremental wear model considering the average pressure has in general a lower slope than the other curves. The difference between the curves from the Wear-Processor and the other wear models discussed is due to the fact that the Wear-Processor considers (i) the continuous decay in the contact pressure as the wear on the pin progresses, (ii) the effect

22.4 Wear Simulation Using the Wear-Processor

Fig. 22-8 Flow chart for the computation using the global incremental wear model: p is the contact pressure, FN is the applied normal load, a is the contact radius due to elastic displacement and wear, ht is the total geometry change at the pin tip, Rp is the curvature of the pin, he is the elastic displacement, hw is the current wear depth, kD is the dimensional wear coefficient, Ds is the interval of the sliding distance, Ec is the elastic modulus of the equivalent surface, smax is the maximum sliding distance and i is the current wear increment number

of friction and (iii) the wear on the counter surface. Even though the global incremental wear models yield a non-linear curve, the simplifying assumptions made in these models are, therefore, considerable. The computation of the maximum pressure in the contact as 1.5 times higher than the average pressure may only be applicable in the initial stage of sliding when the contact is Hertzian. However, for a large sliding distance we assume that the pressure in the contact should approach a ‘flattened’ distribution. Hence the global incremental wear model considering the average pressure represents a lower boundary of the FEM-based solution while the global wear model considering the initial peak pressure is the upper boundary. Between them, the global incremental wear model considering the maximum pressure provides a good approximation for the running-in behavior, as can be seen from its comparatively good agreement with the curve obtained from the Wear-Processor. However, it should be noted that the global wear models described so far have the limitation that they apply only to simple geometries of the interacting surfaces, apart from the other drawbacks discussed earlier. For any general geometry, e.g., the wear on the

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gear flanks of the SFB 499’s demonstrator, we need a local wear model that considers the local time-dependent pressure field, as is the case with the Wear-Processor.

22.5 Conclusion

In this chapter we have given a brief review of the research and advances in the field of modeling and simulation of wear for both macro- and microscale applications. The Wear-Processor described here implements a linear wear law. As an example, the Wear-Processor was used to simulate wear in a pin-on-disc problem, where the pin initially has a spherical tip. The results from the wear simulation compare favorably with the experimental results qualitatively. The accuracy of the FE-based simulation of dry sliding wear is very sensitive to the size of the wear step and also on the FE discretization of the contacting bodies. The Wear-Processor computes the wear on both interacting surfaces and employs an efficient re-meshing technique. For the purpose of further development of the currently linear wear model, additional phenomena, such as pressure dependence and velocity dependence of the wear coefficient, can be easily implemented. Then such a software tool provides a useful platform to investigate such phenomena separately and, moreover, the effects of the incorporated parameters and non-linearities that are of significance to wear. One of the future strategies could be to compare the final topology of the evolved surfaces in the experiments with that from the simulation. In such a way, we would be able to see if additional parameters are necessary or to tune the existing parameters in order to fit the results from the simulation to the experiments. The long-term aim is to use the Wear-Processor in an extended form to predict the progress of wear and the life span of a micromachine.

22.6 Acknowledgment

The authors thank the German Research Foundation (DFG) for funding this work within the scope of the collaborative research center SFB 499 ‘Design, Production and Quality Assurance of Molded Microparts Constructed from Metals and Ceramics’.

22.7 References

22.7 References 1 K. J. Gabriel, F. Behi, R. Mahadevan, 2 3

4 5

6 7 8

9 10 11 12 13 14 15 16 17 18 19 20

21 22 23

24

Sens. Actuators A 1990, 21–23, 184–188. J. A. Williams, Wear 2001, 251, 965–972. K.-H. Zum Gahr, Material- und Verfahrensentwicklung für mikrotechnische Hochleistungsbauteile; FZKA 6662, Karlsruhe: Forschungszentrum Karlsruhe, 2001. N. Huber, J. Aktaa, Microsyst. Technol. 2003, 9, 465–469. K.-H. Zum Gahr, Microstructure and Wear of Materials; Amsterdam: Elsevier Science, 1987, pp. 351–495. J. A. Williams, Wear 1999, 225–229, 1–17. H. C. Meng, K. C. Ludema, Wear 1999, 181–183, 443–457. H. C. Meng, Wear modeling: evaluation and categorization of wear models. PhD Thesis, University of Michigan, Ann Arbor, MI, 1994. B. N. J. Persson, Sliding Friction; Heidelberg: Springer, 2000. S. M. Hsu, M. C. Shen, A. W. Ruff, Tribol. Int. 1997, 30, 377–383. P. J. Blau, Tribol. Int. 1997, 30, 321–331. J. F. Archard, J. Appl. Phys. 1953, 24, 981–988. D. A. Rigney, Wear 1994, 175, 63–69. J. K. Lancaster, Proc. R. Soc. London, Ser. A 1962, 273, 466–483. N. P. Suh, Wear 1973, 25, 111–124. R. Liu, D. Y. Li, Wear 2001, 251, 956– 964. N. Vishwanath, D. G. Bellow, Wear 1995, 181–183, 42–49. A. K. Mukhopadhay, A. Chanda, J. Mater. Sci. Lett. 1999, 18, 149–151. R. Kumar, B. Prakash, A. Sethuramiah, Wear 2002, 252, 445–453. J. F. Molinari, M. Ortiz, R. Radovitzky, E. A. Repetto, Eng. Comput. 2001, 18, 592–609. A. Cantizano, A. Carnicero, G. Zavarise, Comput. Mater. Sci. 2002, 25, 54–60. S. C. Lim, M. F. Ashby, Acta Metall. 1987, 35, 1–24. C. Agelet de Saracibar, M. Chiumenti, Comput. Methods Appl. Mech. Eng. 1999, 177, 401–426. W. Yan, N. P. O’Dowd, E. P. Busso, J. Mech. Phys. Sol. 2002, 50, 449–470.

25 N. Stalin-Muller, K. Van Dang, Wear

1997, 203–204, 180–186. 26 C. Christofides, P. E. McHugh, A. Forn,

27 28 29

30 31 32 33 34 35 36 37 38

39 40 41

42

43 44 45 46 47

J. A. Picas, Comput. Mater. Sci. 2002, 25, 61–72. A. Kapoor, K. L. Johnson, Proc. R. Soc. London, Ser. A 1994, 445, 367–381. M. Öquist, Wear 2001, 249, 6–11. P. Podra, FE wear simulation of sliding contacts. PhD Thesis, Royal Institute of Technology (KTH), Stockholm, 1997. P. Podra, S. Andersson, Tribol. Int. 1999, 32, 71–81. D.-C. Ko, D.-H. Kim, B.-M. Kim, Wear 2002, 252, 859–869. D. J. Dickrell, W. G. Sawyer, Tribol. Trans. 2004, 47, 257–262. W. G. Sawyer, Tribol. Lett. 2004, 17, 139– 145. P. Podra, S. Andersson, Wear 1997, 207, 79–85. Y. Zhao, L. Chang, Wear 2002, 252, 220– 226. W. G. Sawyer, Tribol. Lett. 2004, 17, 139– 145. B. Bhushan, J. N. Israelachvili, U. Lamndman, Nature 1995, 374, 607–616. ABAQUS/Standard 6.4 Users Manual, Vols. I, II, III; Pawtucket, RI: Hibbit, Karlsson and Soresen, USA, 2003. V. Hegadekatte, N. Huber, O. Kraft, Modell. Simul. Mater. Sci. Eng. 2005, 13, 57–75. C. Mattheck, Design in Nature; Heidelberg: Springer, 1998, p. 33. D. M. Tanner, J. A. Walraven, L. W. Irwin, M. T. Dugger, N. F. Smith, W. P. Eaton, W. M. Miller, S. L. Miller, in: IEEE Int. Rel. Phys. Symp., March 1999; 1999, pp. 189–197. K. L. Johnson, Contact Mechanics; Cambridge: Cambridge University Press, 1985, pp. 84–106. H. So, Tribol. Int. 1996, 29, 415–483. H. Goto, Y. Amamoto, Wear 2003, 254, 1256–1266. K. M. Khedkar, J. Khanna, A. S. Gupt, Wear 1997, 205, 220–227. J. Jiang, R. D. Arnell, Wear 1998, 217, 190–199. W. C. Oliver, G. M. Pharr, J. Mater. Res. 1992, 7, 1564–1583.

623

XVII

Subject Index a ABAQUS 36, 120, 561, 610 ablation – laser-based removal 131 ff – mask projection 143 – mold inserts 98 – PIM automation 460 abrasion – binders 293 – mold inserts 93, 97 – surface conditioning 240 f – tribologics 579, 598–604 – wear 609 absorbers 192 accuracy – EDM 99, 170 – feedstocks 300 – industrial robots 479 – LIGA 189 – microassemblies 475–498 – microcasting 357 f, 385 – mold inserts 89 acrylamide/acrylate 433 actuators 7, 476 additives 294 adhesion 328 adhesion grippers 456 ff, 490 agarose 434 age hardening 370, 387, 524 agglomeration 328, 332 air trapping 347, 484 f, 563 alkylphenol polyglycol ether 345 alloys – electroforming 395–395 – microcasting 369 ff – micro-metal injection molding 292, 296

alumina – ceramic replication 423 ff, 427 ff – electrophoretic deposition 430 – gel casting 434 – steel milling 113 – tribologics 579–604 aluminum oxide – EDM 175 – investment casting 374 – lCIM 342 ammonia 375 ammonium iodide 176 ammonium phosphate 375 amorphous nickel phosphorus alloys 90 amorphous polymers 257 annealing 401, 544 ff application specific design 22 arc flame melting 364 Archard wear model 605–624 Armco-iron 113, 246 arrays – electroforming 406 – electrophoretic deposition 430 – microgrinding 97 – PZT 424 Arrhenius type dependence 145 aspect ratio – EDM 99, 164, 170 – electroforming 395 ff, 400, 408 – laser etching 149, 153 – LIGA 189 – lost molds 275 – LPIM 344 – microcasting 357, 388 – lCIM 328 – micro-metal injection molding 322

Advanced Micro and Nanosystems Vol. 3. Microengineering of Metals and Ceramics. Edited by H. Baltes, O. Brand, G. K. Fedder, C. Hierold, J. Korvink, O. Tabata Copyright © 2005 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim ISBN: 3-527-31208-0

XVIII

Subject Index – mold inserts 89 ff, 93, 97 – quality assurance 501 – tribologics 581 assembly concepts 475 ff ATN (Al2O3/TiN) 342 atomic force microscopy (AFM) attenuators 142 austenite 485 austenizing 223, 232 f automation 449–521 AZ9260 resist 192

504, 605 ff

b backlash 483, 489 Bagley correction 76 bainitic materials 238 batch handling automation 453 batch production 88, 477 beam scanning 135 beeswax 294 bending tests – HPIM 341 – microcasting 381 – micro-metal injection molding 316 – micro-PIM 81 – modeling 30, 36 – Stabilor G® 526 ff, 547 ff – steel milling 112 – zirconia 555, 560–578 beryllium wafers 192, 213 binary systems 371 binders – ceramic replication 422 ff – HPIM 331 f, 339 – investment casting 374 – LPIM 345 – lCIM 327 ff – microinjection molding 271, 303 – micro-metal injection molding 290 ff, 293 ff, 298 – micro-PIM 71, 271, 453, 463 – tribologics 581 – zirconia 556 Bingham flow 302 biocompatibility 369 biopolymers 434 Biosil f 373–394 blasting 240 f, 359 blending 290 blowhole-like defects 276 boiling points, etchants 145, 207 Boltzmann constant 58 boric acid 397

boring 102, 291 borosilicon carbonitride 439 boundary conditions 23 ff – handling 466 – lCIM 343 – micro-PIM 68 Bragg peaks 559 brass – diamond cutting 90 – EDM 166 – microinjection molding 268 – tribologics 580 ff, 586 ff, 597 ff bridging-the-gap 53 brittleness – lCIM 326 – steel milling 113 – surface conditioning 224 – zirconia 571 bronze 357–394 brown compacts 453, 467 bubbles – cavitation 245 – electrophoretic deposition 429 – HPIM 339 – zirconia 563 built-up edges 225 bulk temperature 76 bulk viscosity 54 burrs – EDM 179 – steel milling 125 – surface conditioning 221, 229 ff, 236 butyldiglycol (BDG) 199

c calcium sulfate dihydrate 377 calibration 44, 516 capability, steel milling 125 capillaries 56, 175 capillary forces 490 capillary micromolding 421 ff carbide tools 90 ff, 109 carbides – surface conditioning 223, 234 – tribologics 581 see also: cemented carbides carbon absorption 397 carbon black 271,404, 409 carbon fiber-filled plastics 271 carboxylic acid 331 casting 357–394 – alloys 369 ff – ceramic replication 421 ff

Subject Index – design flow 18 – lCIM 326 – Stabilor G® 523–554 catalytic decomposition 310 Catamold – HPIM 330 – micro-metal injection molding 294 – tribologics 583 ff – zirconia feedstocks 558 cavities – EDM 98 – electroforming 403 f, 411 – LBM 102 – LIGA 203 – microinjection molding 254, 259 f, 270 – micro-metal injection molding 290, 304, 309, 322 – micro-PIM 54, 70 – mold inserts 87 ff – tribologics 580 – wet peening 245 cemented carbides – diamond cutting 90 – electrodes 173, 178 – feedstocks 301 – micro-metal injection molding 289, 316 – steel milling 109, 114 – surface conditioning 224, 234 f – tribologics 581 centrifugal casting 360, 367 f, 386 – Stabilor G® 523–554 ceramic components 273 ceramic crucibles 364 ceramic feedstocks – HPIM 336 – micro-metal injection molding 292 – tribologics 579 ff ceramic injection molding (CIM) 325–356 ceramic powders 271 ceramic process technology 253 ff ceramic reactor system 354 ceramic replication techniques 421–448 ceramic slurries 373 ceramic suspensions 441 CFX-4 package 54 chain linked polymers 55 chamfer width 121 Chapman–Enskog method 72 characterization – mold inserts 579–604 – Stabilor G® 528 ff – zirconia 562 ff

chemical composition, Stabilor G® 527 chemical degradation 294 chemical deposition see: electroless depostion chemical etching 133 f chemical-mechanical polishing 610 chemical properties, gold base materials 369 chip thickness 227 chitosan 434 chromatographic electroforming 397 chrome-quartz masks, l 141, 192 clamping 254 ff, 262 f, 304 cleaning 207, 397, 414 climb milling 126 C-Mold package 55 coagulation casting 421, 436 ff coatings – investment casting 379 – LIGA 193, 213 – mold inserts 93 – tribologics 581 cobalt-based hard metals 110 cobalt powders 300 CoCrMo alloy 357 ff, 372–394 codeposition 402 coil angle 121 cold ablation 143 colloids 53, 422, 434 columns 189 compactors 298 compliant mechanism 486 ff compounding feedstocks – HPIM 331 – LPIM 346 – MIM 298 ff compression injection molding 256, 278 compression strength 112 compressive testing 224, 229 computational fluid dynamics (CFD) 52 ff computer aided design (CAD) 12, 21 ff, 53 concentricity errors 223 conductivity – electroforming 395, 401, 408 – LIGA 194 – mask projection 143 – microinjection molding 259 – micro-PIM 76, 80 contact angles 70 contact areas, wear 606, 611 ff, 618 ff contaminants 607 continuous furnace 471

XIX

XX

Subject Index continuum-based computational fluid dynamics (c-CFD) 57 ff continuum theory 52 ff convex polyhedrons 31 cooling – microcasting 383 – microinjection molding 259 – micro-metal injection molding 291, 312 – PIM automation 470 – Stabilor G® 523, 528 ff – zirconia 557 f coordinate measuring machine (CMM) 210, 512 copper – diamond cutting 90 – EDM 100, 166 – electroforming 403, 416 – feedstocks 301 – laser-based removal 138 – layers 194, 202 – Stabilor G® 524 copper alloys 318, 371 copper sulfuric-type plating 202 correlation image, Stabilor G® 536 corrosion, laser etching 148 corrosion resistant castings 371 ff counter structures 132 CPM9V, tribologics 559–604 cracks – HPIM 339 – LIGA 209 – lCIM 329 – micro-metal injection molding 294 – surface conditioning 231 cristobalite 374, 377 Cross model 75 crosslinked polymer solvent gel casting 433 f crucibles 364 crystal growth 365 crystalline melting temperatures 271 current correlation function 73 cutting – EDM 165 – mold inserts 88 ff, 95 – PIM automation 460 – steel milling 109, 123 – surface conditioning 223 ff, 230 ff cuvette embedding 377 cycles 64, 308 cyclic three-point bending, Stabilor G® 549 f

d dark etching layers 232 dark-field image, Stabilor G® 546 debinding – HPIM 338 – lCIM 329 – micro-metal injection molding 289 ff, 310 ff – PIM 454, 468 ff – zirconia 557 deburring 222, 241 Debye–Hückel equation 428 decomposition – HPIM 339 – investment casting 376 – mask projection 143 – microcasting 370 – microinjection molding 271 – micro-metal injection molding 310 – zirconia 564 defects – LIGA 209 – microcasting 365 – microinjection molding 259, 271, 276 – modeling 31 – zirconia 564, 572 deflasking 376 deflection – industrial robots 489 – LIGA 204 – steel milling 115 f, 123 – zirconia 574 deformation – demolding 268 – LPIM 347 – Stabilor G® 525 ff, 549 ff – surface conditioning 229, 239 – wear 607 degassing, zirconia 562 degradation – electroforming 402 – HPIM 339 – LIGA 203, 213 – microinjection molding 260 – micro-metal injection molding 294, 308 ff degrees of freedom (DOF), industrial robots 481 Deguvest 379 f, 385 f, 527, 534 f demagnifying mask projection 143 demolding 261, 461, 464 f dendrites 365, 372 dental casting 357–394

Subject Index dental gold alloy, Stabilor G® 523–554 dental prothesis surfaces 241 depths – engraving 103 – laser etching 149 – LIGA 210 design 1–85, 3–28, 454 see also: system design design-rule-check (DRC) 21, 24 ff destruction tribologics 580 developer solution 194, 199 dewetting 68 diamond tools 90 f, 109, 269 dibenzol peroxide (DBPO) 195 dicing 95 die casting 358 die cavities 290 die sinking EDM (SEDM) 100, 161–174 dielectric fluids 163 Diesel effect 259, 308 differential scanning calorimetry (DSC) 76 diffraction – LIGA 209 ff – Stabilor G® 545 – zirconia 559 diffusion interface 402 diffusity 76 diffusor components 81 dilatometers 78, 303 dimensional measurement technology 499–520 dimensional product qualification 385, 406 dimensions, manufacturing 17 dimethylaniline (DMA) 195 DIN EN ISO 4288 529, 558 diode pumped Nd:YAG laser 102 Dirac d function 58 direct coagulation casting (DCC) 421, 436 ff direct writing, laser-based removal 131, 135 ff direction change 80 discrete components 478 dislocation movement 525 dispersion – ceramic replication 422 – electroforming 400 – HPIM 331 – LPIM 345 – lCIM 329 dissipative particle dynamics (DPD) 51–62

dissociation 427 dissolution – EDM 163 – laser etching 148 – micro-metal injection molding distortions 294, 580 disulfitoaureate 399 doctor blade process 422 domain growth 53 DoseSim 196 double layers 428 drilling – EDM 100, 165 – LBM 102 – micro-metal injection molding – mold inserts 88 ff, 93 – quality assurance 510 – tribologics 581 drop molds 367 dry micropeening 222, 247 drying 346, 442 ductility 233, 400 dynamics, microproduction 7

294

291

e edge radius, zirconia 555, 561–578 effect analysis, quality assurance 500 ejection systems – microinjection molding 262, 266 – micro-metal injection molding 305 – PIM automation 460 elastic anisotropy 29–50 elastoplastic deformation 229 electric microinjection molding 262 electrical conductivity 194 electrochemical deposition 395 ff electrochemical replication 405 electrochemical surface conditioning 240 electrodes – ceramic replication 421, 430 f – EDM 163, 168 – mold inserts 100 electrodischarge machining (EDM) 161–190 – LIGA 205 – mold inserts 88 ff, 98 – steel milling 108, 117 f – tribologics 581 – zirconia 559 electroforming – LIGA 201 – lost molds 275 – metals 395–420

XXI

XXII

Subject Index – microinjection molding 268 ff – mold inserts 187 ff electrohydraulic effect (EHE) 164 electroless metal deposition 90, 403 electrolytes 204, 276, 397 ff electromagnetic effect (EME) 164 electromagnetic radiation 101 electronspectroscopy for chemical analysis (ESCA) 414 electrophoretic deposition (EPD) 421, 426 ff electroplating – accuracy 476 – automation 452 ff – ceramic particles 402 – laser-induced removal 152 ff – LIGA 194 – lost molds 275 electrostatic grippers 456 ff, 490 ff electrosteric stabilization 427 f electrotactile sensors 506 elongation-to-fracture – microcasting 370 – Stabilor G® 531 – surface conditioning 229 embodiment design unit 20 embossing 421–427, 442 EN ISO 9283, industrial robots 479, 489 end milling 107–130, 222, 248 endeffector 479 engraving 102, 165 environment, design 3–28 enzyme catalyzed reactions 436 EPON SU-8 resists 194, 215, 406, 417 epoxy-based resists 214, 443 equation-of-state (EOS) 59 f equipment technology 478 erosion 416, 579 ff etching – accuracy 476 – laser-based 131 ff, 144 ff, 149 – layers 232 – microinjection molding 268 ethyl acetate 207 Euler angles 32 eutectic point 372 evacuation techniques 259 ff excimer laser 102, 140, 152 expansion, thermal 76 f extruders 296, 332, 346

f fabrication – automation 451 ff – electrophoretic deposition 429 ff – lithographic 187–219 – micro-metal injection molding 290 f – quality assurance 508 – soft molding 441 failure – LIGA 213 – lCIM 343 – steel milling 118 – wear 606, 609 – zirconia 566, 572 f failure modes and effect analysis (FMEA) 500 fatigue resistance 112 fault-tree analysis (FTA) 500 feeding, industrial robots 479 feedstocks – HPIM 330 – LPIM 345 – lCIM 325–356 – microinjection molding 87, 261 ff, 271, 273 f – micro-metal injection molding 290 ff, 293 ff – micro-PIM 71 f, 460, 463 – quality assurance 509 f – soft molding 438 – tribologics 579–604 – zirconia 556 FEPAKOM (production of passive optical components) 502 ferrites 98, 223 ferrous metals 110 ferrules 256 fiber probes 389, 505 ff field distortions 136 fillers – investment casting 374 ff – microcasting 388 – microinjection molding 260 f, 264 f finishing 205 f, 240 finite element method (FEM) 29, 33, 36, 57 – Stabilor G® 526 – wear 605–624 – zirconia 556, 561–578 flame retardants 581 flank angle 115 ff, 121 f flexure hinges 475, 483 ff, 487 ff

Subject Index flow characteristics – HPIM 335 – microcasting 388 – micro-PIM 52–59, 67 ff flow chart see: system design FLOW-3D 54 fluorescence 212 fly cutting 91 flying mask 152 focused ion beam (FIB) 115 focusing laser-based removal 134 ff foil turning 91 Fokker–Planck–Boltzmann equation 72 form errors 99, 229 form filling 366, 388, 403, 411 fracture surfaces – Stabilor G® 531, 537, 540 f, 551 – zirconia 560, 566 ff fracture toughness 112, 224, 229 ff fragile microparts 455, 504 freezing 259 frequencies 7, 134 Fresnel diffraction 90, 97, 211 friction – industrial robots 483, 489 – micro-PIM 56 f, 69, 72 f – mold inserts 579–604 – wear 607 ff frozen-wall particles 69 furnaces 471

g Galilean invariance 57 galvanic deburring 241 galvanic forming 397 – laser-induced removal 154 – seed layer 201 – Stabilor G® 529 galvanometer scanning 135 gas sensor set-up 443 gate systems 266 Gauss distribution 68 Gaussian modes 133 ff, 144, 152 G-CuAl10Ni alloy 371 gearhousing 172 gearwheels – gel casting 434 – HPIM 337 – LPIM 354 – micro-metal injection molding 306 – modeling 42 – soft molding 441 – steel milling 127

gel casting 421, 432 ff geometric effects – design flow 10, 23 ff – EDM 170 – microcasting 389 – microcomponents 47 – micro-PIM 73, 80 – steel milling 121, 127 – tool edges 90, 95 glass fibers 256 glass-transition temperature 65, 271 gluing 195 gold absorbers 192 gold-based alloys – copper 205, 525 – dental 31–48 – microcasting 357–394 – silver copper 370 – Stabilor G® 523–554 gold layers 202, 205 gold sulfite electrolytes 399 gold sulfuric plating 202 grain boundary hardening 525 grain dissolution 163 grain growth 289, 296, 328 grain orientation 29, 41–50 grain size – microcasting 383 – micro-metal injection molding 314, 317 – Stabilor G® 533 ff – tribologics 583 – zirconia 556, 564 ff grain structures 31, 112 granulates 259 f, 264 f, 453 graphite – EDM electrodes 100 – investment casting 375 – microcasting 364 grating structures 189, 202 gravitational forces 264, 351 gray-scale correlation image, Stabilor G® 536 green compacts 302 ff – ceramic replication 422 ff – gel casting 432 – lCIM 326 f, 343 ff – PIM automation 453, 467 grinding – EDM 161 – microinjection molding 276 – mold inserts 88 ff – PIM automation 460

XXIII

XXIV

Subject Index – steel milling 109 – surface conditioning 224 – tribologics 581 – wet peening 246 grippers – industrial robots 479 – microassemblies 475, 490 ff – PIM automation 455 ff, 459 ff, 465 ff grooves – EDM 99 – laser etching 147 ff – laser-based removal 132 ff – microgrinding 97 – microinjection molding 266 – steel milling 109, 119 – surface conditioning 222, 229, 234 – Sysmelec Autoplace 411 492 – WEDG 178 gyration radius 54

h hair, microgrinded 98 Hall–Petch relation 525, 543, 546 f handling 451, 465 ff, 490 ff hard metals 289, 310 ff Hardcore Poisson process 32 hardened steels – end milling 107–130 – laser etching 148 – microinjection molding 270 – surface conditioning 221 hardening – laser-based removal 138 – microcasting 370 – Stabilor G® 524 f – surface conditioning 231, 240 f hardness – LIGA 210 ff – nickel alloys 401 – microcasting 370 – mold inserts 579–604 – steel milling 112, 122 – surface conditioning 224, 232 ff – wear 608 HASYLAB beamline 558 heat capacity 59, 76 heat exchanger plate 425 heat flow 79 heat treatments 221–234, 289 heating – laser-based removal 134 – microcasting 363 – micro-metal injection molding 312

– PIM automation 460 Hele–Shaw flow model 54 Helmholtz double layers 397 Helmholtz–Smucholski equation 428 Henry factor 428 Hertz formula 615 ff, 620 ff high precision 195, 512 high-pressure capillary rheometer (HPCR) 302, 334, 341 high-pressure injection molding (HPIM) 325, 330 ff – zirconia 555–578 high-speed tool steels (HSS) 93 high-strength materials 372 ff hole-type templates 412 holes 99, 177 homogenizer 142 hot-crack susceptibility 364 hot embossing – laser-based removal 132 ff, 152, 155 – LIGA 188 – mask projection 143 hot isostatic pressing 312 hot molding 344 hot pressing 100 housing – EDM 172 – hermetically sealed 154 – laser-induced removal 154 – microcasting 382 – microinjection molding 257, 27, 29, 307 hybrid parallel robots MICABO 483 hydraulic microinjection molding 262 hydrocarbon chains 331 hydrogen athmosphere 318 hydrogen evolution 397 hydroxyapatite (HAP) 423 ff hysteresis curves 148

i IGES design flow 21 illumination 145 image structuring 143 impurities 397 incrementals 620 induction heating 364 industrial environments 170 industrial robots 478 ff inert electrodes 399 infiltrating 379 information units 20 injection molding 93

Subject Index injection molding – laser-based removal 132 ff – mask projection 143 – PIM automation 451–474 – quality assurance 502 – tribologics 579–604 – zirconia 555–578 injection pressure 327 injection transfer molding (ITM) 279 injection velocity 308, 341 injector components 81 in-mold assembly 253, 274 f insert injection molding 274 f, 305, 409 inserts see: mold inserts insulating walls 275 integrated circuits (ICs) 4 intermediate masks 192, 205 investment casting 357 ff, 368 f, 373 ff, 386 ion beaming 114 ion etching 268 ion-exchange cleaning 397 ionization 427 iridium segregation 525 iron/alloys 271, 296 see also: steel irradiation 145, 190, 199 island-type templates 412 ISO standards, industrial robots 479, 489 isoelectric point (IEP) 437 isometric drawing 443 isotropic behavior 31 isotropic finishing 241

j jet chemical etching 146 ff jewelry casting 357–394 jewelry housings 291 joints 277

k Karush–Kuhn–Tucker condition 44 kinematic behavior, industrial robots 480, 485 kinematic viscosity 73 kneaders 290 f, 33, 346 knowledge-based engineering (KBE) 3, 24 ff KoMeth database 21, 24 f

l labeling 102 Langevin approach 53 laser ablation

– EDM 161 – steel milling 116 – tribologics 581 laser-based removal 131–160 laser-based surface conditioning 240 laser-beam machining (LBM) 88 ff, 101 – design flow 6 – steel milling 115 laser etching 144 ff laser-flash method 76 laser patterning 270 lathe machine 179 layers – electroforming 403 – electrophoretic deposition 428, 432 – LIGA 190 ff, 201 ff, 213 – microcasting 371 – micro-metal injection molding 314 – surface conditioning 232 lead zirconate titanate (PZT) – ceramic replication 423 ff, 432 ff – gel casting 436 – soft molding 439 Lennard–Jones potential 61 lenses 209, 136 Licomont® 294, 330, 334 life times 500 LIGA (Lithografie Galvanoformung Abformung) 4, 187 ff – ceramic replication 423 ff – electroforming 399, 404 – laser-based removal 132 – microcasting 359 – microinjection molding 255, 268 – mold inserts 88, 100 – Stabilor G® 524, 429 – worldwide facilities 216 ligaments 170 light microscopy 528, 558 line widths 170 linear elastic materials 31 linear regression 565 liquid-phase jet chemical etching 144 ff liquid silicone rubber (LSR) 259 liquid vapor systems 53 ff, 61 lithium aluminate structure 424 lithography – LBM 102 – microinjection molding 268 – mold inserts 187–219 – soft molding 437 – steel milling 107 loading, zirconia 555

XXV

XXVI

Subject Index look-and-move procedures 496 lost molds 275, 357, 360 ff low pressure injection molding (LPIM) 325 ff, 344 ff – zirconia 555–578 low viscosity feedstocks 438 Lowe–Andersen thermostat 58 ff, 63, 72, 80 lubrication 579–607

m machining – accuracy 476 – LPIM 348 f – microinjection molding 262 – micro-metal injection molding 304 – surface conditioning 221 magnesium pyrophosphate 375 manufacturing – automation 451 ff – construction requirements 507 – design flow 13 ff – microassemblies 491 – lCIM 326 ff – mold inserts 87–106 – polymer microparts 413 – steel milling 114 f many-body dissipative particle dynamics (MDPD) 60 f Martens hardness 225 martensite 485 martensitic structures 223 mask projection method 140 ff masking – electroforming 395 – LIGA 188 – substrate package 196 material-flow automation 467 material properties 521–624 – mold inserts 221–254 – tribological 579–604 materials – microproduction 7 – steel milling 111 Maxwell distribution 60, 68, 72 measurement technology 499–520 mechanical erosion techniques 416 mechanical feedstock data 261 mechanical grippers 456 f, 466 f, 490 f mechanical microproduction 5 ff mechanical properties – Biosil 373 f – gold base alloys 370

– microcasting 381, 387 – Stabilor G® 523–554 – surface conditioning 222 ff, 230, 234 – zirconia 555–578 melting – investment casting 375 – laser-based removal 134 – microcasting 363 – microinjection molding 254 melting points – LIGA 212 – microinjection molding 271 ff – micro-metal injection molding 294, 317 – paraffin 345 membrane methods 421 ff mesastructures 132 f mesh generation 33 mesoscopic colloid dynamics 53 mesoscopic heat flow 79 metal electroforming 395–420 metal growth 408 metal injection molding (MIM) 289–330 metal-ion complexation 434 metallic feedstocks 336 metallic microbending 34 metallization 213, 414 metallographics 253, 312 metallurgical alloys 292 methacrylamide (MAM) 434 methological design flow 3 ff methylcellulose 379 methylenebisacrylamide (MBAM) 433 methylethyl ketone (MEK) 424 methylmethacrylate (MMA) 195 MICABO parallel robots 481 ff, 494 ff microassemblies 475–498 microbending tests 30, 36 – HPIM 341 – Stabilor G® 526 ff, 547 ff microcasting 357–394 – design flow 18 – quality assurance 513 – Stabilor G® 523–554 microcavities – microinjection molding 270 – micro-metal injection molding 304, 309 – micro-PIM 54 – tribologics 580 microceramic injection molding (lCIM) 325–356 microchannels 54 microcomponents – electroforming 400

Subject Index – microinjection molding 255 – modeling 35, 42 – properties 521–624 – tribologics 579–604 microcutting 108 microdrilling 93, 268, 581 microelectrodischarge machining (lEDM) 161–186 – microinjection molding 268 – zirconia 559 see also: electrodischarge machining microelectroforming 395–420 microelectromechanical systems (MEMS) 98, 606 microend milling 107–130, 222 microfeatures 499, 508 ff microgears 89, 441, 276, 279, 512 MicroGlider 386, 406 microgrinding 95 microgrippers 455 468, 490 ff microholes 95 microhousing 271 microinjection molding 108, 152 ff, 253–292 micromachining 451 ff, 605–624 micromechanical components 453, 508 micro-metal injection molding 289–330 micrometrology, dimensional 499–520 micromilling – casting 360 – design flow 6, 13 – injection molding 86, 268 – surface conditioning 221, 240 – tribologics 581 – zirconia 556 micromolding-in-capillaries (MIMIC) 421, 438 ff micromolds 107, 161 see also: molds, molding microoptical systems (MOEMS) 189 microparts – casting 381 – CIM 325–356 – injection molding 253 ff micropatterning – ceramic replication 422 – microinjection molding 255, 266 ff – soft molding 442 micropeening 221, 240 f micropowder injection molding (lPIM) 51–86, 253, 289 – design flow 6, 14 ff – quality assurance 502 ff, 513 ff

– tribologics 579–604 microprecision components 190, 255 microproduction 5 ff micropunching 180 f microreplication techniques 437 microrings 337 microsegregation 365 microspecific design flow 7 microspecimen 34 – Stabilor G® 526 – zirconia 555–578 microstructures 231 – components 421, 429 ff – electroforming 402 – microcasting 357 f, 383, 390 – micro-metal injection molding 289, 313 – modeling 29–50 – Stabilor G® 523–554 – tribologics 586 ff – zirconia 555–578 microsystem technology (MST) 254, 397 microtechnology – design flow 4 ff – lCIM 325–356 – wear 605–624 microtensile testing 36 ff – HPIM 340 – Stabilor G® 526–538 ff microthree-point bending, zirconia 555, 561–578 microtransfer molding (lTM) 421, 438 ff microtribology 579–604 microturbine 353 microwave plasma treatment 414 milling – design flow 6, 13 – EDM 161, 167 ff, 172 ff, 177 – microcasting 360 – microinjection molding 258 – micro-metal injection molding 291 – mold inserts 92 – quality assurance 510 – surface conditioning 221, 238 – tribologics 581 – zirconia 556 miniaturization – automation 451 ff – mold inserts 88 – steel milling 107 minimum sizes 16 ff mirrors 209 Mises stress distribution 38 f, 44, 118 mixing phenomenons 59

XXVII

XXVIII

Subject Index mobility, electrophoretic 428 modeling 29–50, 605–624 mold inserts – design flow 13 – EDM 172 – electroforming 402 – fabrication 87–106, 187 ff – investment casting 379 – laser based removal 131–160 – microinjection molding 266 ff – micro-metal injection molding 305, 309 – micro-PIM 70 – specifications 87–106, 187 ff – steel milling 127 – tribologics 579–604 MoldFlow simulation 55, 261 molding – automation 451–474 – ceramic replication 421 ff – high-pressure 336 – LIGA 188 – materials 271 – quality assurance 502, 513 – steel milling 108 – zirconia 555–578 see also: micro-, injection, etc. molecular dynamics 53 molochite 374 molybdenum wires 166 momentum vectors 57 monocrystalline alumina 113 monocrystalline silicon 95 moving masks 141 MPM 9 steel 122 MSC.Patran meshing 33 mullite 374 multicomponent microinjection molding 81 multilayer compounded components 477 Multivac 368

n NanoFocus 386, 558 Navier–Stokes equation 57 Nd:YAG lasers 102, 132 ff Newtonian fluids 54, 75, 347 nickel – electroforming 203, 395–420 – feedstocks 301 – microinjection molding 268 – tribologics 580, 597 nickel alloys – diamond cutting 90

– electroforming 395–420 – laser etching 144 ff – microinjection molding 268 – micro-metal injection molding 292, 318 nickel–gold interface 205 nickel layers 201 f nickel sulfamate plating 202 nickel surfaces 139 no-slip condition 67 nodal displacement 44 noise amplitude 58 noncontact laser-based methods 134 nondirected nickel electroplating 403 ff nonsilicon cadidates 188 notch factor 40, 575 nozzles – drilling 511 – EDM 172 – HPIM 337, 341 – laser etching 144 – microcasting 382 nucleation 365

o object/operation system 3 ff octadecylamine 345 open flame technique 363 optical glass fibers 256 optical probing 502 ff, 505 ff optimization 44 ff optotactile sensors 512 ordering hardening 525 outsertmolding 274 overmolding 274, 304 Owens–Wendt–Rabel–Kaelble method 414 oxide layers 371

p PA12-C 271 packaging 462, 477 f pale etching layers 232 paraffin – electrophoretic deposition 430 – LPIM 345 – lCIM 327 – zirconia 557 parallel assembly 477 ff parallel robots 475, 480 particle packing theory 430 particles 154, 290, 297, 328 passivation 147 passive optical components 502 PATRAN 33, 612

Subject Index patterning – investment casting 379 – LIGA 212 – LPIM 344 – microcasting 362 – microinjection molding 254, 266 ff – soft molding 442 payload 483 ff PCBs 96 peak-to-valley roughness 533 ff, 569 ff PEEK 271, 275, 410 peening 221, 240 f permittivity 428 pH-catalyzed coagulation 436 pH values 203 phase transitions 59, 357 phosphate-bonded investments 368, 375 phosphoric acid 144 photoablation 133 f photochemical decomposition 143 photoelectrons 211 photolithography 476 photometric electroforming 397 photoresist casting (PRC) 442 physical vapor deposition (PVD) 201 pin-on-disk wear tests 605 ff piston-type machines 348 planar robots 481, 486 plasma-enhanced silicon etching 270 plaster-bonded investments 377 ff, 534 plastic deformation 233, 239 – lCIM 329 – Stabilor G® 523 ff, 549 ff plastic flow 425 plastic injection molding 455 plastics 253, 259 platicizers 308, 422 plating 202, 397 platinum 145, 399 pneumatic demolding 461 pneumatic grippers 466 pneumatic machines 348 pneumatic tactile sensor 506 Poisson process 32 Poisson ratio – Stabilor G® 531 – steel milling 120 – zirconia 562 polarographic electroforming 397 polishing – laser-based removal 138 – LIGA 196, 206 – microcasting 359

– steel milling 109 – surface conditioning 240 – wear 610 poly(methyl methacrylate) (PMMA) – ceramic replication 423 ff – electroforming 403, 410, 413 f – investment casting 380 – LIGA 187, 192 f, 200 – microcasting 361 – microinjection molding 257 ff, 271, 275 poly(vinyl alcohol) 331, 434 polyamide 331, 410 polycarbonate – laser-induced removal 152 – mask projection 143 – microinjection molding 256 ff, 271 polycrystals 31, 526, 606 polydimethylsiloxane (PDMS) 437, 441 polyethoxymethylsiloxane 439 polyethylene – HPIM 331 – microinjection molding 257 ff, 271 – micro-metal injection molding 294 polyfin wax 330 polyglycol ethers 345 polymer binders 290 polymer ceramic composites 266 polymer molds 397 polymer precursors 438 polymer surfaces 395 polymeric microcomponents 580 polymerization 195, 432 polymers – laser-based removal 133 f – LIGA 188 – mask projection 143 – lCIM 328 f – microinjection molding 253, 257 ff, 264 ff – micro-PIM 54 polyoxymethylene (POM) – electroforming 410 – HPIM 330 – microinjection molding 257 ff, 271, 275, 278 – micro-metal injection molding 294 polypropylene 294, 331, 398 polyststyrene 55, 410 polysulfone (PSU) 257, 271 polyvinylalcohol (PVA) 294 pore-containing microtensile testing 39 porosity – microcasting 364

XXIX

XXX

Subject Index – micro-metal injection molding 313 – Stabilor G® 533 ff – zirconia 564 ff, 570 ff positioning – EDM 99, 170 – microassembling 491, 497 – PIM automation 465 postprocessor, wear 609 posttreatments 416 potassium bromide/chloride/iodide 176 potter’s clay minerals 374 powder binder compounds – HPIM 331 – micro-metal injection molding 298 – zirconia 556 powder binder separation 303 powder injection molding (PIM) 31 ff, 51–86 – automation 451–474 – ceramic 328 f – low-pressure 344 ff – metal 289, 296 – tribologics 579–604 preblending 290 preceramic polymers 439 precipitation – electroforming 398 – hardening 525 – microcasting 370 – surface conditioning 223 precision components 190, 195 precision engineering – casting 359 f – laser-based removal 132 – micro-metal injection molding 291 – steel milling 108 – structure elements 463 precision limiting effects 211 f precursors 432 ff, 438 pressure management 308 – capillary rheometer 75 – hot isostatic pressing 312 – microcasting 386 – lCIM 326 – microinjection molding 260 – micro-PIM 64, 76 Prestomat 368 primary structuring 395, 402 principle see: system design process chains – 2C injection molding 410 – design flow 6, 24 f – LIGA 191

– microinjection molding 263 – mold fabrication 152 ff – PIM automation 462 – specific design 22 process parameters – electroforming 397 – micro-metal injection molding 304, 308 f – micro-PIM 51 – steel milling 125 product life-time 500 production 5 ff, 11, 24 f – HPIM 340 – passive optical components (FEPAKOM) 502 – quality assurance 500 – Stabilor G® 526 ff – zirconia specimen 556 projection techniques 134, 140 ff, 152 propane oxygen flame 363 protective layers 432 proteins 434 proximity lithography 211 pseudoelastic flexure hinges 483 ff pulse-power technologies 164 pulsed laser systems 101, 133 punching 180 f purity requirements 397 pyrolysis 360, 432 PZT see: lead zirconate titanate

q qualilty function deployment (QFD) 500 quality assurance 449–521 – dimensual measurements 499–520 – electroforming 402 – LIGA 187 quarter model 561 quartz 374, 377 quasi-isotropic behavior, Stabilor G® 526 quasi-static loading 555 quasi-static micro-three-point bending 548 quenching – microcasting 371, 384 – Stabilor G® 525 – surface conditioning 233, 242

r radiation 187 rake angle 115 ff, 121 f random heat flux 79 rapid processing 135

Subject Index raw binder materials 295 Rayleigh distribution 68 reactive ion etching (RIE) 268 reflecting boundary conditions 68 reflectivity 143, 209 refraction 209 refractory fillers 377 rehydration 377 remelting 138 removal technologies 131–160 repeatability 479, 482 ff, 489 ff repetition rates 135 replication techniques 5, 13 ff, 251–448 – ceramic 421–448 – EDM 162, 202 – electroforming 395, 416 – laser-induced removal 152 – LIGA 188 – microcasting 385, 390 – mold inserts 87, 131 ff – molding 438 – polymer molds 402, 405 ff – zirconia 556 repulsion coefficients 63 residual organic components 199 residual stresses 555–578 residues 209 resin-bonded dicing blades 97 resist layers 190 ff, 193 ff, 406 resist substrate mask 196, 204, 211 resistance heating 363 resistivity 112, 175 resolidification process 138 resolution – industrial robots 479 – laser etching 144 – laser-based removal 134 f – measurement technology 504 – nickel electroplating 405 retightening 45 Reynolds number 56 rheological behavior – feedstocks 300 ff – HPIM 333, 341 – LPIM 347 – micro-PIM 53 Ribcon connector 257 robots 478 ff Rollomatic software 120 rotary joints 482 ff rough machining 138, 173 roughness – EDM 99

– – – – – – – –

laser-induced removal 153 LIGA 189, 210 microinjection molding 270 Stabilor G® 523–554 steel milling 108 surface conditioning 221, 235 f wear 608 zirconia 569 f see also: surface roughness RPS-AH parallel robot 481 runner systems 266 rupture – LPIM 344 – microcasting 370 – steel milling 115

s sand blasting 359 sandwich technique 196 sapphire cutting tools 113 saturated calomel electrode (SCE) 145 scanning electron microscopy (SEM) – alumina 586 ff – tribologics 586 ff – X175VcrMo9–5 591 f – zirconia 558–588 scanning methods – laser-based removal 135 f – quality assurance 503 f f – structuring 137 Schmid number 72 scratch test 605–624 scribing 97 secondary radiation 187 secondary structuring 395 seed layer 201 segregation – microcasting 364 – microinjection molding 260 – micro-PIM 59, 463 – Stabilor G® 525, 532 self-isolating effect 259 semicrystalline polymers 257 semihydrates 377 sensor guidance 475, 483 ff, 491 ff separation automation 458 ff serial assembly 476 serial robots 480 shape memory alloy (SMA) 485 f shaping methods – ceramic replication 426 – laser etching 149 – LIGA 189

XXXI

XXXII

Subject Index – lCIM 326 ff – optimization 45 shear rates 76 shear roll compactors 298 shear thinning 302, 347 shear viscosity 73 f, 557 shock resistance 112, 163 shrinkage – ceramic replication 423 – electrophoretic deposition 430 – feedstocks 298 – HPIM 337 – microcasting 364 – microcomponents 45 – micro-metal injection molding 305, 319 – quality assurance 513 – soft molding 442 side milling 223, 227 sidewall roughness 189 sigma-blade kneader 290, 346 silica 376, 380 silicon – microinjection molding 268, 279 – microsystems 4 – tribologics 580 – wafers 204 – wear 606 – zirconia 563 silicon carbide 435 silicon nitride 423 ff, 435 silicon technology – automation 452 – etching 270 – membranes 506 – mold inserts 88, 98 silicone rubber – electroforming 416 – LPIM 350 – microinjection molding 259 – zirconia 556 Siliplast 294, 330 siloxane polyglycol ether 345 silver 524 simulation – microcomponents 35 f – microinjection molding 261 – micro-PIM 51–86 – steel milling 118 – wear 605–624 see also: finite element method single-crystals 31, 526 single-flute end milling 115 single-grain materials 32

sintering – ceramic replication 423 ff – electrophoretic deposition 430 ff – HPIM 338 f – microcasting 360 – lCIM 329 – micro-metal injection molding 289 ff, 300–322 – PIM automation 462, 467 ff – zirconia 557 size effect, zirconia 555, 564 ff sliding contact – tribologics 579–604 – wear 608, 613 ff slurries 373, 422 smooth finishing 240 sodium chloride 176 soft molding 421, 437 ff soft tooling 350 sol–gel electrophoretic deposition 432 ff solidification – investment casting 374 – microcasting 357, 365, 369 – microinjection molding 254 – soft molding 438 solvent-assisted micromolding (SAMIM) 421, 438 ff solvent evaporation 310, 422 spark erosion technique 416 sparks 163, 176 spatial robots 481, 487 specific surface area 334 specific volume 64 specifications – design flow 10, 24 – LIGA 210 ff – mold inserts 187 – robots 478 ff spinneret tool 170 spinning 200 spinodal decomposition 370 spring–damper system 55 sprue systems 266, 453, 460 f sputtering 414 stability 409, 427 stabilized zirconia 556 Stabilor G® 523–554 – microcasting 361–392 – PIM 31 ff, 48 stacking 109, 462, 467 stainless steels 296, 322, 349 standard deviation 43 f standard processes, LIGA 191

Subject Index standardization 7 statistical process control (SPC) 500 stearic acid 294 steel – end milling 107–130 – LPIM 349 – microdrilling 96 – microinjection molding 270, 289–312 – mold inserts 91 ff – surface conditioning 233 – tribologics 579–604 steel 17–4PH 31–48, 310, 325 steel 30CrMo6, tribologics 582 ff steel C80 588 steel MPM 9 122 steel Toolox 44 122, 233, 582 ff steel wires 166 steel X175VCrMo9–5 579–604 steel X38CrMoV5–1 – milling 115, 118, 124 – tribologics 579–604 STEP design flow 21, 24 ff stick-slip effect 483 stiffness 43 stochastic differential equation 53 strain energy 37 stress distribution 29–50 stress strain curves – industrial robots 485 ff – Stabilor G® 534 ff, 541 ff, 546 ff stress tensor 611, 616 ff stress values 40 ff – microcasting 370, 387 – zirconia 560 ff striations 240, 244 f stripping 207, 397 structural optimization 44 structures – EDM 99 – electrophoretic deposition 431 – LIGA 189, 198, 210 – microcasting 388 – PIM automation 463 – steel milling 127 structuring – electroforming 395 – laser-based removal 132 ff – LBM 102 f stylus profilometry 138 SU-8 resins 189, 192 f, 214 sub-angström resolution 504 substrates – ceramic replication 422

– electroforming 395 – LIGA 193, 204 sulfamate 276, 397, 403 sulfate 397 sulfur release 377 sulfuric acid 144 supercooling 532 superhard plasters 377 surface-close material states 239 surface conditioning 221–254 surface fluid interactions 54 surface forces 490 surface heating 134 surface images – Stabilor G® 536 – tribologics 588 ff, 595 ff surface layers 314 surface mounted devices (SMD) 476 ff surface node 611 ff, 617 ff surface quality – conditioning 227 – laser-based removal 131 ff – LIGA 196 – LPIM 351 – micro-metal injection molding 296 – mold inserts 89 ff, 93 ff – PIM automation 463 – steel milling 108, 123 surface roughness – ceramic replication 426 – investment casting 378 – laser etching 151 ff – laser-based removal 137 – microcasting 357, 386 f – lCIM 328 f – mold inserts 89 ff, 93 ff – steel milling 123 – vacuum pressure microcasting 367 – wear 608 surface-to-volume ratio 6 – HPIM 330 – microinjection molding 259 f, 264 f – micro-metal injection molding 295 – wear 605 ff – zirconia 556, 565 f surface topography – Stabilor G® 523–554 – zirconia 555–578 surface walls 54 surfactants – electroforming 397 – HPIM 331 – LIGA 203

XXXIII

XXXIV

Subject Index – LPIM 345 surgical implants 372 susceptibility 364 suspensions – electroforming 402 – electrophoretic deposition 427 – soft molding 441 swelling 187 synchronized scanning 141 synchrotron radiation 558 Sysmelec Autoplace 411 480, 492 ff system design – centrifugal microcasting 367 – direct writing 136 – EDM 163 – electroforming 404 – electrophoretic deposition 427 ff – gel casting 433 – grippers 466 – laser etching 144 – LIGA 191, 197 – LPIM 349 – mask projection 142 – microassemblies 478, 495 – microcasting 361 – microinjection molding 254 – micro-metal injection molding 291 – micropart separation 460 – mold fabrication 152 – PIM automation 464 – surface conditioning 233 – triboinjection molding 584 – vacuum pressure microcasting 366 – WearProcessor 612 – WEDG 166, 181 – zirconia 559

t tactile methods 505 ff, 512 f Tait parameters 64 f tape casting 421 ff target system definition 3, 8 f TEA CO2 laser 102 technology specification sheets 10, 24 temperature–time profiles 310, 339 temperatures, 64 – LIGA 202, 213 – microcasting 363, 386 – microinjection molding 267, 271 – micro-metal injection molding 308 – Stabilor G® 532 – surface conditioning 223, 236 – zirconia 557

tempering 233, 238, 242, 266 f templates 191, 408, 414 ff tensile testing – HPIM 340, 343 – microcasting 272, 362, 370, 381, 387 – microinjection molding 272 – micro-metal injection molding 316, 322 – Stabilor G® 524–538 ff – surface conditioning 224, 234 ternary AuAgCu alloys 370 test geometries 80, 127 testing methods 30, 36, 556 ff textures – LPIM 351 – modeling 32 – surface conditioning 221, 244 ff thermal conductivity 143, 259 thermal data, feedstocks 261 thermal decomposition 564 thermal degradation 294, 308, 310 ff thermal effects 254 – laser-based removal 133 – investment casting 374 – LIGA 212 – micro-PIM 76 f thermal loading 211 thermal shock 112, 163 thermal treatments – HPIM 338 ff – LPIM 347 ff – micro-metal injection molding 310 ff – surface softening 226 thermogravimetric analysis (TGA) 351 thermoplast 341 thermoplastic materials – ceramic replication 425 – lCIM 326 – microinjection molding 260, 276 – micro-metal injection molding 294 thermoplasticity 609 thin films 422 three-axis micromilling 15 three-dimensional microstructuring 131– 160 three-dimensional molded interconnect device (3D-MID) 277 three-linear drives of the structure see: TRIGLIDE three point bending 30, 36 – Stabilor G® 548 ff – zirconia 555, 561–578 TiAlN coatings 112 Ticast 368

Subject Index time-resolved measurements 145 titanium electrodes 399 titanium membranes 192 titanium nitride 342 titanium sapphire laser 102 tolerances 195 f tooling 85–250 – electrodes 163 – geometries 225 f – LPIM 348 f – micro-metal injection molding 304 – milling 111, 107 ff – PIM automation 458 ff – steels 233 ff, 579–604 Toolox 44® 122, 233, 582 ff topographic measurements 406 torques 298, 333, 581 TOSCA.shape optimization 44 toughness 222, 227 trace metal impurities 397 transducers 424 transfer molding see: injection transfer molding 278 transformation points 138 transforming processes 108 transition temperatures 65, 374 transmission electron microscopy (TEM) 528, 532, 544 transverse current correlation function 73 tree pattern 362 tribological properties – micro-metal injection molding 316 – mold inserts 579–604 – surface conditioning 233 f TRIGLIDE spatial parallel robot 481 tube dilatometer 78 tungsten 100, 301, 316 ff turbine – gel casting 435 – LPIM 353 – microcasting 382 – micro-metal injection molding 307 twin-screw extruder 298, 346 two-component injection molding 253, 276, 319, 342 two-component templates 411 two-finger grippers 466 two-step replication 416

u ultimate tensile strength 370, 387 ultimate tensile stress 530, 539

ultrafine grains 112 ultrasonic-assisted diamond turning process 110 ultrasonic-supported stripping 207 ultrasonic transducers 424 ultrasonic wet peening 245 UMAP-130 probe 505 uncertainty 514 ff unigraphics 24 UV lithography 214 ff

v vacuum grippers 456 ff, 490 ff vacuum pressure casting (VPC) 360, 365 f, 386, 523–554 validation 7, 29–50 van der Waals fluids 62, 490 vaporizing 134 variance of multiple direction position accuracy (vAP) 480 ff, 489 ff Variotherm processs 259 ff, 266 ff, 308 Venturi principle 466 Verlet algorithm 58 very large-scale integrated circuits (VLSI) 5 Vickers hardness 225, 588 f Victor–Kienzle equation 123 virtual testing 30 viscosity, binders 294 viscosity – electrophoretic deposition 428 – feedstocks 298, 303 – LPIM 345 ff – lCIM 327 ff – micro-PIM 54, 72 – soft molding 438 – tribologics 599 – zirconia 334, 557, 564 ff visual sensor guidance 483 ff voids 569 ff, 576 f voltammetric electroforming 397 volume effect – investment casting 378 – micro-PIM 64 – zirconia 566 von Mises stress distribution 38 f, 44, 118 Voronoi tessellation 31 ff

w wafers 192, 204 wall thickness – HPIM 339 – microcasting 370

XXXV

XXXVI

Subject Index – microinjection molding 260 – micro-PIM 55, 70 Warren’s approach 63 Washburn equation 56, 67 water content, investment casting 375 water lubrication 579–604 Watts-type electrolytes 397 waxes 302 – HPIM 330, 341 – LPIM 345 – microcasting 360, 368 – lCIM 327 ff – micro-metal injection molding 294 WC–Co electrodes 173 wear – electrodes 168 – micro-metal injection molding 304 – mold inserts 579–604 – simulation 605–624 – steel milling 127 – surface conditioning 222 wear-resistant materials 91 ff, 165 WearProcessor 605–624 webs 99, 109 Weibull theory – gel casting 435 – lCIM 328 – zirconia 555, 560–570 ff weight functions 58 Weissenberg–Rabinowitsch correction 76 welding 155 Werth fiber probe (WFP) 210, 505 wet chemical processes 204, 268 wet peening 222, 245 wetting 59, 68 f, 369 wheels 382 white-light microscope 558 wicking process 310, 351 wire electrodischarge grinding (WEDG) 100, 161 f, 165 ff Wöhler diagram 550 working temperatures see: temperatures

workpieces – EDM 99, 163 – steel milling 122 ff – surface conditioning 229 workspace, industrial robots 482 ff, 488 ff workstation, direct writing 136 worst case assumption 47

x X-ray lithography – ceramic replication 425 – electroforming 399 – LIGA 211, 187 ff X38CrMoV5–1 steel – milling 115, 118, 124 – surface conditioning 233 f, 242

y yield point 345 ff yield strength 120 yield stress 370, 387, 525 Young’s equation 70 Young’s modulus – gold 526 – Stabilor G® 530, 539 – steel milling 112, 120 – tribologics 583 – zirconia 562 yttria stabilized zirconia ceramics 582

z zirconia – ceramic replication 423 ff, 427 ff – HPIM 333 f, 337 – LPIM 347, 353 – lCIM 328 – micro-PIM 78 – PIM 31 ff, 36 ff, 42 ff, 48 – tribologics 579–604 zirconia-toughened alumina (ZTA) 435 zirconium oxide 374 zirconium silicate 374 z-kneaders 290