Modern Techniques of Coating and Drying in Surface Engineering 1774071991, 9781774071991

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MODERN TECHNIQUES OF COATING AND DRYING IN SURFACE ENGINEERING

MODERN TECHNIQUES OF COATING AND DRYING IN SURFACE ENGINEERING

Edited by:

Maria Emilova Velinova

ARCLER

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www.arclerpress.com

Modern Techniques of Coating and Drying in Surface Engineering Maria Emilova Velinova

Arcler Press 2010 Winston Park Drive, 2nd Floor Oakville, ON L6H 5R7 Canada www.arclerpress.com Tel: 001-289-291-7705 001-905-616-2116 Fax: 001-289-291-7601 Email: [email protected] e-book Edition 2020 ISBN: 978-1-77407-384-1 (e-book)

This book contains information obtained from highly regarded resources. Reprinted material sources are indicated. Copyright for individual articles remains with the authors as indicated and published under Creative Commons License. A Wide variety of references are listed. Reasonable efforts have been made to publish reliable data and views articulated in the chapters are those of the individual contributors, and not necessarily those of the editors or publishers. Editors or publishers are not responsible for the accuracy of the information in the published chapters or consequences of their use. The publisher assumes no responsibility for any damage or grievance to the persons or property arising out of the use of any materials, instructions, methods or thoughts in the book. The editors and the publisher have attempted to trace the copyright holders of all material reproduced in this publication and apologize to copyright holders if permission has not been obtained. If any copyright holder has not been acknowledged, please write to us so we may rectify. Notice: Registered trademark of products or corporate names are used only for explanation and identification without intent of infringement. © 2020 Arcler Press ISBN: 978-1-77407-199-1 (Hardcover)

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DECLARATION Some content or chapters in this book are open access copyright free published research work, which is published under Creative Commons License and are indicated with the citation. We are thankful to the publishers and authors of the content and chapters as without them this book wouldn’t have been possible.

ABOUT THE EDITOR

Maria Velinova is Ph.D. holder in Quantum chemistry at the University of Sofia since April 2012. Her major research experience is in the field of Computational Chemistry, especially in statistical mechanics methods applied to different sorts of biomolecules. Member of the Laboratory of Quantum and Computational Chemistry at the University of Sofia.

TABLE OF CONTENTS

List of Contributors .......................................................................................xv List of Abbreviations ................................................................................... xxv Preface.................................................................................................. ....xxix SECTION I: INTRODUCTION TO COATING AND DRYING PROCESSES Chapter 1

Numerical Simulation of the Thin Film Coating Flow in Two-Dimension .. 3 Abstract ..................................................................................................... 3 Introduction ............................................................................................... 4 Coating Fluid Simulation Theory ................................................................ 5 The Result of Calculation and Simulation................................................... 8 Conclusion .............................................................................................. 13 Acknowledgements ................................................................................. 13 References ............................................................................................... 14

Chapter 2

Corrosion Study of Powder-Coated Galvanised Steel .............................. 15 Abstract ................................................................................................... 15 Introduction ............................................................................................. 16 Experimental Details ................................................................................ 17 Results and Discussion ............................................................................ 18 Conclusion .............................................................................................. 27 Acknowledgment ..................................................................................... 27 References ............................................................................................... 28

Chapter 3

Toward a Nearly Defect-free Coating via High-energy Plasma Sparks .... 31 Abstract ................................................................................................... 31 Introduction ............................................................................................. 32 Results and Discussion ............................................................................ 34 Conclusions ............................................................................................. 44

Materials and Methods ............................................................................ 44 References ............................................................................................... 46 SECTION II: COATING AND SURFACE DEFECTS Chapter 4

Carbon based DLC films: Influence of the Processing Parameters on the Structure and Properties ............................................ 53 Abstract ................................................................................................... 53 Introduction ............................................................................................. 54 Materials and Methods ............................................................................ 55 Results and Discussion ............................................................................ 57 Conclusions ............................................................................................. 63 Acknowledgments ................................................................................... 63 References ............................................................................................... 64

Chapter 5

Superiority of Graphene over Polymer Coatings for Prevention of Microbially Induced Corrosion ........................................................... 65 Abstract ................................................................................................... 66 Introduction ............................................................................................. 66 Results and Discussions ........................................................................... 68 Conclusion .............................................................................................. 79 Methods .................................................................................................. 80 Acknowledgements ................................................................................. 83 References ............................................................................................... 84

Chapter 6

Seed Defective Reduction in Automotive Electro-Deposition Coating Process of Truck Cabin .............................................................. 89 Abstract ................................................................................................... 89 Introduction ............................................................................................. 90 Define Phase ........................................................................................... 90 Measure Phase......................................................................................... 91 Analyze Phase ......................................................................................... 93 Improvement Phase ................................................................................. 94 Control Phase .......................................................................................... 95 Conclusion .............................................................................................. 96 References ............................................................................................... 97

x

SECTION III: FEED PREPARATION FOR COATING Chapter 7

Indentations on Air Plasma Sprayed Thermal Barrier Coatings Prepared by Different Starting Granules ............................................... 101 Abstract ................................................................................................. 101 Introduction ........................................................................................... 102 Materials and Experimental Procedures ................................................. 103 Results and Discussion .......................................................................... 107 Conclusion ............................................................................................ 113 Acknowledgments ................................................................................. 114 References ............................................................................................. 115

Chapter 8

In Situ Carbon Coated LiNi0.5Mn1.5O4 Cathode Material Prepared by Prepolymer of Melamine Formaldehyde Resin Assisted Method.................................................................................................. 119 Abstract ................................................................................................. 120 Introduction ........................................................................................... 120 Experimental Sections ............................................................................ 120 Results and Discussion .......................................................................... 122 Conclusions ........................................................................................... 126 Acknowledgments ................................................................................. 126 References ............................................................................................. 127

Chapter 9

Washcoat Deposition of Ni- and Co-ZrO2 Low Surface Area Powders onto Ceramic Open-Cell Foams: Influence of Slurry Formulation and Rheology ...................................... 131 Abstract ................................................................................................. 131 Introduction ........................................................................................... 132 Results And Discussion .......................................................................... 135 Experimental Section ............................................................................. 144 Conclusions ........................................................................................... 147 Acknowledgments ................................................................................. 148 Author Contributions ............................................................................. 148 References ............................................................................................. 149

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SECTION IV: ROLL COATING AND RELATED PROCESSES Chapter 10 Experiment Research on Hot-Rolling Processing of Nonsmooth Pit Surface ............................................................................................. 155 Abstract ................................................................................................. 155 Introduction ........................................................................................... 156 Modeling of Pipe Inner-Wall Machining Robot ...................................... 158 Analysis of Hot-Rolling Machining Robot .............................................. 161 Text Set And Test Method ....................................................................... 166 Test Results And Analysis ....................................................................... 170 Conclusions ........................................................................................... 175 Acknowledgment ................................................................................... 176 References ............................................................................................. 177 Chapter 11 Bioinspired Superhydrophobic Surfaces, Fabricated Through Simple and Scalable Roll-To-Roll Processing ......................................... 181 Abstract ................................................................................................. 181 Introduction ........................................................................................... 182 Results ................................................................................................... 183 Discussion ............................................................................................. 192 Methods ................................................................................................ 192 Acknowledgements ............................................................................... 194 References ............................................................................................. 195 Chapter 12 Roll-to-Roll Slot-die Coating of 400 mm Wide, Flexible, Transparent Ag Nanowire Films for Flexible Touch Screen Panels ........ 199 Abstract ................................................................................................. 199 Introduction ........................................................................................... 200 Results ................................................................................................... 202 Conclusion ............................................................................................ 214 Methods ................................................................................................ 215 Acknowledgements ............................................................................... 217 References ............................................................................................. 218 Chapter 13 Surface Tension-Driven Self-Alignment ................................................. 223 Abstract ................................................................................................. 223 Introduction ........................................................................................... 224

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Fundamentals Of Surface-Tension-Driven Self-Alignment ...................... 227 Liquid Confinement ............................................................................... 236 Influence of Shape and Size ................................................................... 239 Dynamics .............................................................................................. 245 Process Integration................................................................................. 250 Patents ................................................................................................... 264 Outlook ................................................................................................. 265 Acknowledgements ............................................................................... 266 References ............................................................................................. 267 Chapter 14 Protruding Organic Surfaces Triggered by in-plane Electric Fields ........ 279 Abstract ................................................................................................. 280 Introduction ........................................................................................... 280 Results ................................................................................................... 281 Discussion ............................................................................................. 290 Methods ................................................................................................ 290 Acknowledgements ............................................................................... 292 References ............................................................................................. 293 Chapter 15 Quality of Electroless Ni-P (Nickel-Phosphorus) Coatings Applied in Oil Production Equipment with Salinity ............................... 295 Abstract ................................................................................................. 295 Introduction ........................................................................................... 296 Deposition Process of Nickel-Phosphorus .............................................. 297 Mechanism And Properties of Ni-P Deposition ...................................... 297 Ni-P Coating: Specifications And Properties ........................................... 302 Corrosion Testing of Electroless Ni-P Coating ......................................... 307 Concerns Or Precautions Regarding Nickel-Phosphorus Coating Applied on Carbon Steel and Used In Extreme Conditions ........... 307 Conclusions ........................................................................................... 309 References ............................................................................................. 310 SECTION V: DRYING PROCESSES Chapter 16 Laser Alloying Advantages by Dry Coating Metallic Powder Mixtures with SiOx Nanoparticles ........................................................ 315 Abstract ................................................................................................. 316

xiii

Introduction ........................................................................................... 316 Materials and Methods .......................................................................... 320 Results and Discussion .......................................................................... 327 Conclusions ........................................................................................... 347 Patents ................................................................................................... 347 Author Contributions ............................................................................. 348 Acknowledgments ................................................................................. 348 Appendix A ........................................................................................... 348 References ............................................................................................. 350 Chapter 17 Low-Temperature, Dry Transfer-Printing of a Patterned Graphene Monolayer ............................................................................ 361 Abstract ................................................................................................. 361 Introduction ........................................................................................... 362 Transfer-Printing of a Patterned Graphene Layer ..................................... 364 Results and Discussion .......................................................................... 366 Conclusion ............................................................................................ 373 Methods ................................................................................................ 374 Acknowledgements ............................................................................... 376 References ............................................................................................. 377 Index ..................................................................................................... 381

xiv

LIST OF CONTRIBUTORS Zhiyong Liang College of Science, Donghua University, Shanghai, China Hui Zhou College of Science, Donghua University, Shanghai, China Manish Kumar Bhadu R&D and Scientific Services, TATA Steel Ltd., Jamshedpur 831001, India Akshya Kumar Guin R&D and Scientific Services, TATA Steel Ltd., Jamshedpur 831001, India Veena Singh R&D and Scientific Services, TATA Steel Ltd., Jamshedpur 831001, India Shyam K. Choudhary R&D and Scientific Services, TATA Steel Ltd., Jamshedpur 831001, India Mosab Kaseem Materials Electrochemistry Laboratory, School of Materials Science and Engineering, Yeungnam University, Gyeongsan, 38541, Republic of Korea Hae Woong Yang Materials Electrochemistry Laboratory, School of Materials Science and Engineering, Yeungnam University, Gyeongsan, 38541, Republic of Korea Young Gun Ko Materials Electrochemistry Laboratory, School of Materials Science and Engineering, Yeungnam University, Gyeongsan, 38541, Republic of Korea Francisco Andrés Delfín Surface Engineering Group (GIS) – UTN FRCU – Ing. Pereyra 676, Concepción del Uruguay, Argentina. CP: 3260 xv

Sonia Patricia Brühl Surface Engineering Group (GIS) – UTN FRCU – Ing. Pereyra 676, Concepción del Uruguay, Argentina. CP: 3260 Christian Forsich Materials Department – FH-OOE, Wels Campus – Stelzhamerstrasse 23, Wels, Austria. CP: 4600 Daniel Heim Materials Department – FH-OOE, Wels Campus – Stelzhamerstrasse 23, Wels, Austria. CP: 4600 Ajay Krishnamurthy Mechanical, Aerospace and Nuclear Engineering, Rensselaer Polytechnic Institute, 110 8th Street, Troy, New York 12180, USA Venkataramana Gadhamshetty Civil and Environmental Engineering, South Dakota School of Mines and Technology, Rapid City, South Dakota 57701, USA Rahul Mukherjee Mechanical, Aerospace and Nuclear Engineering, Rensselaer Polytechnic Institute, 110 8th Street, Troy, New York 12180, USA Bharath Natarajan Department of Materials Sciences and Engineering, Rensselaer Polytechnic Institute, 110 8th Street, Troy, New York 12180, USA Osman Eksik Mechanical, Aerospace and Nuclear Engineering, Rensselaer Polytechnic Institute, 110 8th Street, Troy, New York 12180, USA S. Ali Shojaee Mechanical and Aerospace Engineering, Oklahoma State University, 218 Engineering North, Stillwater, Oklahoma 74078, USA Don A. Lucca Mechanical and Aerospace Engineering, Oklahoma State University, 218 Engineering North, Stillwater, Oklahoma 74078, USA

xvi

Wencai Ren Shenyang National Lab for Materials Science, Institute of Metal Research, Chinese Academy of Sciences, Shenyang 110016, China Hui-Ming Cheng Shenyang National Lab for Materials Science, Institute of Metal Research, Chinese Academy of Sciences, Shenyang 110016, China Nikhil Koratkar Mechanical, Aerospace and Nuclear Engineering, Rensselaer Polytechnic Institute, 110 8th Street, Troy, New York 12180, USA Department of Materials Sciences and Engineering, Rensselaer Polytechnic Institute, 110 8th Street, Troy, New York 12180, USA Aekkalag Sonthilug Department of Industrial Engineering, Chulalongkorn University, Bangkok, Thailand 10330 Parames Chutima Department of Industrial Engineering, Chulalongkorn University, Bangkok, Thailand 10330 Regional Centre for Manufacturing Systems Engineering, Faculty of Engineering Chulalongkorn University, Bangkok, Thailand 10330 Yong Suk Heo School of Mechanical Systems Engineering, Kookmin University, Seoul 136702, Republic of Korea Dong Heon Lee School of Mechanical Systems Engineering, Kookmin University, Seoul 136702, Republic of Korea Yeon-Gil Jung School of Nano and Advanced Materials Engineering, Changwon National University, Changwon 641-773, Republic of Korea Kee Sung Lee School of Mechanical Systems Engineering, Kookmin University, Seoul 136702, Republic of Korea xvii

Wei Yang School of Chemistry and Chemical Engineering, Guangzhou University, 230 Waihuanxi Road, Panyu District, Guangzhou 510006, China Haifeng Dang School of Chemistry and Environmental Engineering, Dongguan University of Technology, Dongguan 523808, China Shengzhou Chen School of Chemistry and Chemical Engineering, Guangzhou University, 230 Waihuanxi Road, Panyu District, Guangzhou 510006, China Hanbo Zou School of Chemistry and Chemical Engineering, Guangzhou University, 230 Waihuanxi Road, Panyu District, Guangzhou 510006, China Zili Liu School of Chemistry and Chemical Engineering, Guangzhou University, 230 Waihuanxi Road, Panyu District, Guangzhou 510006, China JingLin School of Chemistry and Chemical Engineering, Guangzhou University, 230 Waihuanxi Road, Panyu District, Guangzhou 510006, China Weiming Lin School of Chemistry and Chemical Engineering, Guangzhou University, 230 Waihuanxi Road, Panyu District, Guangzhou 510006, China School of Chemistry and Chemical Engineering, South China University of Technology, 381 Wushan Road, Guangzhou 510640, China Riccardo Balzarotti Politecnico di Milano, Dipartimento di Chimica, Materiali e Ingegneria Chimica “G. Natta”, Piazza Leonardo da Vinci 32, 20133 Milano, Italy Mirko Ciurlia Politecnico di Milano, Dipartimento di Chimica, Materiali e Ingegneria Chimica “G. Natta”, Piazza Leonardo da Vinci 32, 20133 Milano, Italy Cinzia Cristiani Politecnico di Milano, Dipartimento di Chimica, Materiali e Ingegneria Chimica xviii

“G. Natta”, Piazza Leonardo da Vinci 32, 20133 Milano, Italy Fabio Paparella Politecnico di Milano, Dipartimento di Chimica, Materiali e Ingegneria Chimica “G. Natta”, Piazza Leonardo da Vinci 32, 20133 Milano, Italy Yun-qing Gu College of Mechanical Engineering, Zhejiang University of Technology, Hangzhou 310014, China Tian-xing Fan College of Mechanical Engineering, Zhejiang University of Technology, Hangzhou 310014, China Jie-gang Mou College of Mechanical Engineering, Zhejiang University of Technology, Hangzhou 310014, China Wei-bo Yu China Aviation Powerplant Research Institute, Zhuzhou 412002, China Gang Zhao College of Mechanical and Electrical Engineering, Harbin Engineering University, Harbin 150001, China Evan Wang Institut National Polytechnique of Grenoble, Joseph Fourier University, Grenoble 38031, France Sung-Hoon Park Department of Mechanical engineering, Soongsil University, 369 Sangdo-ro, Dongjak-gu, Seoul, 156-743, Korea Sangeui Lee Material Research Center, Samsung Advanced Institute of Technology, Yonginsi, Gyeonggi-do, 446-712, Korea David Moreira Department of Mechanical & Aerospace Engineering, University of California, San Diego, La Jolla, CA 92093- 0411, USA xix

Prabhakar R. Bandaru Department of Mechanical & Aerospace Engineering, University of California, San Diego, La Jolla, CA 92093- 0411, USA InTaek Han Material Research Center, Samsung Advanced Institute of Technology, Yonginsi, Gyeonggi-do, 446-712, Korea Dong-JinYun Material Research Center, Samsung Advanced Institute of Technology, Yonginsi, Gyeonggi-do, 446-712, Korea Dong-Ju Kim Kyung Hee University, Department of Advanced Materials Engineering for Information and Electronics, 1 Seocheon, Yongin, Gyeonggi-do 446-701, Republic of Korea Dynamic Korea Technology, R&D Center, 116-60, Sanho-daero, Gumi City, Gyeong-Buk, 39377, Republic of Korea Hae-In Shin Kyung Hee University, Department of Advanced Materials Engineering for Information and Electronics, 1 Seocheon, Yongin, Gyeonggi-do 446-701, Republic of Korea Eun-Hye Ko Kyung Hee University, Department of Advanced Materials Engineering for Information and Electronics, 1 Seocheon, Yongin, Gyeonggi-do 446-701, Republic of Korea Ki-Hyun Kim Samsung Display, OLED R&D Center, Yongin, Gyeonggi-do 446-711, Republic of Korea Tae-Woong Kim Samsung Display, OLED R&D Center, Yongin, Gyeonggi-do 446-711, Republic of Korea Han-Ki Kim Kyung Hee University, Department of Advanced Materials Engineering for

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Information and Electronics, 1 Seocheon, Yongin, Gyeonggi-do 446-701, Republic of Korea Massimo Mastrangeli Physical Intelligence Department, Max Planck Institute for Intelligent Systems, Max Planck ETH Center for Learning Systems, Heisenbergstr. 3, 70569 Stuttgart, Germany Quan Zhou Department of Electrical Engineering and Automation, School of Electrical Engineering, Aalto University, Otaniementie 17, 02150 Espoo, Finland Veikko Sariola Department of Automation Science and Engineering, Tampere University of Technology, Korkeakoulunkatu 3, 33720 Tampere, Finland Pierre Lambert Department of Bio, Electro And Mechanical Systems, E´cole Polytechnique de Bruxelles, Universite´ Libre de Bruxelles, CP 165/56. Avenue F.D. Roosevelt 50, 1050 Brussels, Belgium Danqing Liu SCNU-TUE Joint Lab of Devices Integrated Responsive Materials (DIRM), South China Normal University, No. 378, West Waihuan Road, Guangzhou Higher Education Mega Center, Guangzhou 510006, China Department of Chemical Engineering, Delft University of Technology, Van der Maasweg 9, 2629 HZ Delft, The Netherlands Nicholas B. Tito Department of Applied Physics, Eindhoven University of Technology, Postbus 513, 5600 MB Eindhoven, The Netherlands Institute for Complex Molecular Systems (ICMS), Eindhoven University of Technology, Den Dolech 2, 5612 AZ Eindhoven, The Netherlands Dirk J. Broer Institute for Complex Molecular Systems (ICMS), Eindhoven University of Technology, Den Dolech 2, 5612 AZ Eindhoven, The Netherlands Laboratory of Functional Organic Materials & Devices (SFD), Department of Chemical Engineering & Chemistry, Eindhoven University of Technology, Den Dolech 2, 5612 AZ Eindhoven, The Netherlands xxi

Fernando B. Mainier Escola de Engenharia, Universidade Federal Fluminense (UFF), Niterói, Brazil Maria P. Cindra Fonseca Escola de Engenharia, Universidade Federal Fluminense (UFF), Niterói, Brazil Sérgio S. M. Tavares Escola de Engenharia, Universidade Federal Fluminense (UFF), Niterói, Brazil Juan M. Pardal Escola de Engenharia, Universidade Federal Fluminense (UFF), Niterói, Brazil Michael C. H. Karg Institute of Photonic Technologies (LPT), Friedrich-Alexander-Universität Erlangen-Nürnberg (FAU), Konrad-Zuse-Straße 3/5, 91052 Erlangen, Germany Collaborative Research Center 814-Additive Manufacturing (CRC 814), Friedrich-Alexander-Universität Erlangen-Nürnberg (FAU), Am Weichselgarten 9, 91058 Erlangen-Tennenlohe, Germany Erlangen Graduate School in Advanced Optical Technologies (SAOT), Friedrich-Alexander-Universität Erlangen-Nürnberg (FAU), Paul Gordan Straße 6, 91052 Erlangen, Germany Michael Rasch Institute of Photonic Technologies (LPT), Friedrich-Alexander-Universität Erlangen-Nürnberg (FAU), Konrad-Zuse-Straße 3/5, 91052 Erlangen, Germany Collaborative Research Center 814-Additive Manufacturing (CRC 814), Friedrich-Alexander-Universität Erlangen-Nürnberg (FAU), Am Weichselgarten 9, 91058 Erlangen-Tennenlohe, Germany Erlangen Graduate School in Advanced Optical Technologies (SAOT), Friedrich-Alexander-Universität Erlangen-Nürnberg (FAU), Paul Gordan Straße 6, 91052 Erlangen, Germany Konstantin Schmidt Institute of Photonic Technologies (LPT), Friedrich-Alexander-Universität Erlangen-Nürnberg (FAU), Konrad-Zuse-Straße 3/5, 91052 Erlangen, Germany Sophia A. E. Spitzer Institute of Photonic Technologies (LPT), Friedrich-Alexander-Universität Erlangen-Nürnberg (FAU), Konrad-Zuse-Straße 3/5, 91052 Erlangen, Germany xxii

Till F. Karsten Institute of Photonic Technologies (LPT), Friedrich-Alexander-Universität Erlangen-Nürnberg (FAU), Konrad-Zuse-Straße 3/5, 91052 Erlangen, Germany Daniel Schlaug Institute of Photonic Technologies (LPT), Friedrich-Alexander-Universität Erlangen-Nürnberg (FAU), Konrad-Zuse-Straße 3/5, 91052 Erlangen, Germany Cosmin-Rudolf Biaciu Institute of Photonic Technologies (LPT), Friedrich-Alexander-Universität Erlangen-Nürnberg (FAU), Konrad-Zuse-Straße 3/5, 91052 Erlangen, Germany Andrey I. Gorunov Kazan National Research Technical University named after A.N. Tupolev-KAI, Karl Marx Str. 10, 420111 Kazan, Russia Michael Schmidt Institute of Photonic Technologies (LPT), Friedrich-Alexander-Universität Erlangen-Nürnberg (FAU), Konrad-Zuse-Straße 3/5, 91052 Erlangen, Germany Collaborative Research Center 814-Additive Manufacturing (CRC 814), Friedrich-Alexander-Universität Erlangen-Nürnberg (FAU), Am Weichselgarten 9, 91058 Erlangen-Tennenlohe, Germany Erlangen Graduate School in Advanced Optical Technologies (SAOT), Friedrich-Alexander-Universität Erlangen-Nürnberg (FAU), Paul Gordan Straße 6, 91052 Erlangen, Germany Sugkyun Cha Program in Nano Science and Technology, Graduate School of Convergence Science and Technology, Seoul National University, Seoul 151-742, Republic of Korea Minjeong Cha Program in Nano Science and Technology, Graduate School of Convergence Science and Technology, Seoul National University, Seoul 151-742, Republic of Korea Seojun Lee Program in Nano Science and Technology, Graduate School of Convergence

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Science and Technology, Seoul National University, Seoul 151-742, Republic of Korea Jin Hyoun Kang Department of Chemistry, Seoul National University, Seoul 151-747, Republic of Korea Changsoon Kim Program in Nano Science and Technology, Graduate School of Convergence Science and Technology, Seoul National University, Seoul 151-742, Republic of Korea Advanced Institutes of Convergence Technology, Suwon, Gyeonggi 443-270, Republic of Korea

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LIST OF ABBREVIATIONS AM

Additive Manufacturing

APS

Air plasma sprayed

ACAs

Anisotropic conductive adhesives

AFM

Atomic force microscope

BEOL

Back-end-of-line

BSE

Back Scattered Electron

CMAS

Calcium-magnesium-aluminosilicate

CNT

Carbon nanotube

CVD

Chemical vapor deposition

CSM

Continuous stiffness measurement

CV

Cyclic voltammetry

DOE

Designing of experiment

DSC

Differential scanning calorimetry

DHM

Digital Holographic Microcopy

DMC

Dimethyl carbonate

DFT

Dry film thickness

DMTA

Dynamic Mechanical Thermal Analysis

EAPs

Electroactive polymers

EIS

Electrochemical impedance spectroscopy

ED

Electro-Deposition Coating

EBPVD

Electron beam physical vapor deposition

EDS

Energy dispersive spectrometer

EDX

Energy dispersive spectrometry

EC

Ethylene carbonate

FMEA

Failure Mode and Effect Analysis

FIB

Focused Ion Beam

FE-SEM

Field emission scanning electron microscope

GI

Galvanised

HF

Hydrogen fluoride

ITO

Indium tin oxide

LMD

Laser Metal Deposition

LEDs

Light-emitting diodes

LCNs

Liquid crystal polymer networks

LG

Liquid–gas

LPR

Liquid photo resist

MSE

Mean square error

MSA

Measurement System Analysis

MIC

Microbially induced corrosion

MEMS

Microelectromechanical systems

MWCNTs

Multiwalled carbon nanotubes

OTU

Operational taxonomic unit

PACVD

Plasma-assisted chemical vapor deposition

PSD

Particle Size Distribution

PEO

Plasma electrolytic oxidation

PDMS

Polydimethylsiloxane

PP

Polypropylene

PU

Polyurethane

PVA

Polyvinyl alcohol

KOH

Potassium hydroxide

PCB

Printed circuit board

RIE

Reactive-ion etch

RT

Room temperature

SAM

Self-assembled monolayer

SCE

Saturated calomel electrode

SEM

Scanning electron microscopy

STEM-EDS

Scanning transmission electron microscope capable of energy dispersive x-ray spectroscopy

SA

Sliding angle

SG

Solid–gas

xxvi

SRB

Sulfur reducing bacteria

SH

Superhydrophobic

TBCs

Thermal barrier coatings

TWCs

Three way catalysts

TSPs

Touch screen panels

TEM

Transmission electron microscopy

TCEs

Transparent conductive electrodes

WC

Tungsten carbide

UTS

Ultimate Tensile Strength

UV

Ultra-violet

VIAs

Vertical interconnect accesses

VOC

Volatile organic content

WCA

Water contact angle

XRF

X-ray fluorescence spectroscopy

XPS

X-ray photoelectron spectroscopy

YS

Yield strength

YZA

Yttria-stabilized Zirconia Alumina

YSZ

Yttrium stabilized zirconia

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PREFACE

Film coating and drying processes are at the base of modern surface engineering and are applied in numerous fields because they play an important role in the processing, such as improving the surface properties, fine processing, and new surface properties. The aim of this book is to aid decision making when selecting the most suitable coating methods by giving from a materials-science perspective on overview of available techniques. Section 1 of Modern Techniques of Coating and Drying in Surface Engineering book introduces the film coating and drying processes by discussing several kinds of applications. Among those, after it shows a theoretical-numerical study of the film coating process, it treats of powder-coated galvanized steel, of coating via high-energy plasma, of soap-film coating and graphene/epoxy coating for aircraft structures. In the end, Section 1 reviews coating processes based on Zirconium and/or Titanium. Section 2 focuses on coating processes and surface defects. In particular, it discusses the influence of the processing parameters on the structure and properties of the surface, of polymer coatings for prevention of microbially induced corrosion, of seed defective reduction in the electro-deposition coating process, and of TEM and atom probe tomography imaging analysis. Lastly, it deals with changing od defects structure and properties of nanostructured TiSi-N coatings, before and after annealing. Section 3 treats the feed preparation for coating. In particular, it discusses the usage of granules, resins and foams. Section 4 deals with roll coating and related processes. In detail, it presents an experimental study on hot-rolling processing of nonsmooth pit surface, then it focuses on roll-to-roll processing both for bioinspired superhydrophobic surfaces and flexible touch screen panels. Moreover, it focuses on enhancing the adhesion between the polyethylene separator. Section 5 treats the surface tension driven defects and static electricity. In particular, after it reviews the current knowledge and state-of-the-art of surface tension-driven self-alignment, it describes a method to create dynamic surface topographies by an in-plane electrical field, it presents a study on quality of

electroless Ni-P (Nickel-Phosphorus) coatings applied in oil production. In the end, it discusses the fabrication of polymeric coatings using an electrospraying technique. Finally, the last Section 6 discusses the advantages of dry coating metallic powder mixtures via laser alloying, and the dry transfer-printing of a patterned graphene monolayer.

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SECTION I: INTRODUCTION TO COATING AND DRYING PROCESSES

Chapter 1

Numerical Simulation of the Thin Film Coating Flow in Two-Dimension

Zhiyong Liang, Hui Zhou College of Science, Donghua University, Shanghai, China

ABSTRACT Thin film coating is a process of making liquid film cover and deposit base body surface by the way of dipping, spraying, sliding or spin coating, which is a kind of modern surface engineering. It plays an important role in the actual processing, such as improving the surface properties, fine processing, and new surface properties. Analysis of the influence of substrating morphology and fluid flow properties itself on coating fluid motion has an important significance to optimize the thin film coating and improve the quality of the final film. The influence from uneven substrate surface’s geometry configuration on internal motion of the flow field in slip-coating is analyzed by using the FLUENT software as a calculation platform. A

Citation: Liang, Z. and Zhou, H. (2017) “Numerical Simulation of the Thin Film Coating Flow in Two-Dimension”. Open Journal of Fluid Dynamics, 7, 330-339. https://doi.org/10.4236/ojfd.2017.73021 Copyright © 2017 by authors and Scientific Research Publishing Inc. This work is licensed under the Creative Commons Attribution International License (CC BY 4.0).

4

Modern Techniques of Coating and Drying in Surface Engineering

two-dimension model of slip coating under isosceles triangle and isosceles trapezoid substrate was established, and thin film coating fluid motions under different configuration parameters were simulated. It is pointed out that the key factor determining the turbulence generation and evolution is the parameter of substrating surface nature. The effects of the change of Reynolds number on turbulent appearance and action area are studied. The velocity contours of fluid field on different substrate surfaces are shown, and the impact of substrate geometry on the backwater region is analyzed.

Keywords: Numerical Simulation, Thin Film, Slip-Coating, Surface Engineering, Fluid Mechanics

INTRODUCTION Thin film coating is a kind of modern surface engineering, which generally refers to the process of covering substrate surface with a layer of liquid film. At present, coating technology including electroplating, painting, thermal spraying and vapor deposition, compared with the heat treatment, bead welding and other surface engineering, has several advantages such as less constraint conditions, large space of selection technology and material type and so on, and is used in practical engineering more and more widely. It can play out in three ways such as fine processing, optimization of surface properties, and making new surface properties [1] . Specifically, thin film coating can be used to improve the coated substrate surface rust-proof property, anti-corrosion property, abrasions resistance, heat resistance and other characteristics, and has been applied in many fields. In the machinery industry, bridge engineering and electric power industry, thin film coating technique is used to improve micro damage phenomenon of closely matched components [2] ; In the field of Aeronautics, film coating technology is often used in aviation engine parts to improve engine efficiency to prolong its service life [3] ; In the field of ship protection, application of coating technology can greatly improve the imputrescibility and rust preventing characteristics of shipboard equipment [4] . With the development of material science, liquid thin film begins to take a more and more important role, and consequential quality requirements of the coated thin film are more and more high. In modern industrial applications, thin film coating process needs to meet some special requirements such as: the shape of thin film flow surface is complicated, the fluid coating needs

Numerical Simulation of the Thin Film Coating Flow in Two-Dimension

5

to be carried out in a great disturbance; problem of the actual flow of the film is nonlinear; rheological properties of thin film flow cannot be changed arbitrarily; it should meet the demand of high speed coating industrial production. According to the form of thin film defect, it could be divided into two types, discontinuity and continuity [5] . In the actual production, due to factors of mechanical equipment, operation, production process and fluid properties, coating film maybe have some defects which cannot be completely eliminated such as folds, ripple and bubble [6] [7] [8] . Coating defects will certainly affect the coating quality, thereby affecting the properties of the thin film. As the requirements of coating quality and aesthetic appearance of products continuously increase, how to control or reduce the defects of film coating appearance has become an urgent problem to solve, and this needs to do some numerical simulation in the coating fluid mechanics research. In this paper, research purposes are to analyze the influence of substrating morphology and fluid flow properties itself on coating fluid motion to optimize the thin film coating and improve the quality of the final film.

COATING FLUID SIMULATION THEORY In fluid dynamics foreign scholars have carried out much research work on the film flow. However, in this field, most previous studies about the flow of the film focused on the flat surface problem. In recent years, research of the liquid film flow characteristics on surface of basal with specific geometry morphology begins. The reason for this study is the surface of the thin film coating substrate is never perfectly flat but quite complex in the actual production.

Couette Flow Model The film flow on smooth surface can be approximated as a shear flow between two infinite flat plates with rigid surface and specific geometry morphology, which is called the Couette flow. This is a basic research model of the thin film coating. Two-dimension Couette model is shown in Figure 1. The lower flat plate is fixed, the upper mobile flat plate moves at a constant velocity v along the x direction. The velocity distribution between the two flat plates is as follows:

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Modern Techniques of Coating and Drying in Surface Engineering

(1-1) The boundary conditions are as follows: (1-2) where u is the distribution velocity, μ is fluid viscosity, p is the pressure, h is the distance between the parallel plates, x, y represents the spatial coordinates respectively, and v is a constant velocity along the x direction. The velocity distribution can be generated by substituting Formula (1-2) into Formula (1-1) as follows: (1-3) The velocity distribution of dimensionless form is expressed as (1-4) In the above formula, D is the dimensionless pressure gradient

Calculation Model and the Control Equations Thin film coating fluid mechanics problems studied in this paper are based on the improvement of the Couette flow model, in which the bottom surface is undulate substrate.

Figure 1: Couette flow diagram.

Numerical Simulation of the Thin Film Coating Flow in Two-Dimension

7

The change of the internal structure of the flow field caused by the uneven geometry is studied. Physical models of the shear drag flow by the approximate Couette model are shown in Figure 2(a) and Figure 2(c), and the calculation models see Figure 2(b) and Figure 2(d). Among them, the top of model is a flat plate which can do translation, and the backplane silhouette is a cyclical fluctuations curve with isosceles triangle or isosceles trapezoid groove. As with the Couette model, the lower base plate is fixed, and the upper flat plate can move at a constant velocity v along the x direction. Under the hypothesis that the roof and floor are infinite and the base plate geometry is periodic, one circle can be selected for calculation, as shown in Figure 2. Thin film coating fluid in general can be regarded as the Newton fluid, to satisfy the control equation (1-5) (1-6) (1-7) In the formula, u is the velocity vector, ρ is the density, t is the time, T is the temperature, k is heat conduction coefficient, cp is specific heat capacity, and ST is viscous dissipation. In Figure 2: (a) the physical model with isosceles triangle; (b) the computational model with isosceles triangle; (c) the physical model with isosceles trapezoid groove; (d) the computational model with isosceles trapezoid groove. Usually, Formula (1-5) is called the momentum conservation equation, Formula (1-6) is called the mass conservation equation, and Formula (1-7) is called the energy conservation equation.

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Modern Techniques of Coating and Drying in Surface Engineering

Figure 2: Physical model and computational model of the backplane silhouette is a cyclical fluctuations curve.

Those formulas constitute equations of Newton fluid motion and are also known as the Navier-Stokes equations. Reference to Couette flow model, on the roof to be applied moving wall boundary condition, the floor uses the solid wall boundary condition.

THE RESULT OF CALCULATION AND SIMULATION Influence of Roughness of Triangular Base Plate on the Flow Field From the calculation results, compared with Couette flow on the flat base plate, with a fixed plate distance H = 1.6 and constant initial Reynolds number Re = 10, substrate irregularity degree is a key factor affecting the formation of eddy current. Seen from Figure 3, with the increase of basal plate roughness, flow separation phenomenon of thin film coating fluid occurs, accompanied by the vortices. Increasing unevenness r will make the eddy current phenomenon more and more obvious, and effect the location of vortex generation.

Numerical Simulation of the Thin Film Coating Flow in Two-Dimension

9

Figure 3: Stream lines distribution with different roughness r at identical plate distance H = 1.6 and initial Reynolds number Re = 10 on triangular basal plate. (a) r = 1/6; (b) r = 0.3; (c) r = 0.625; (d) r = 1.

Here defined unflatness of the basal plate with triangular groove is r=(H−h)/W .

Influence of Roughness of Trapezoid Substrate on the Flow Field Thin-film coating flow on trapezoidal base board and triangular basal floor are analogous. Similar with the approximate Couette flow on triangular basal plate, as seen from Figure 4, with the increase of basal plate roughness r, flow separation phenomenon of thin film coating fluid field occurs, accompanied with gradual appearing of eddy current. After eddy arising, the increasing of roughness will make the vortex phenomenon more and more obvious. In other words, increasing the roughness of R will expand the scope of the vortex.

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Modern Techniques of Coating and Drying in Surface Engineering

Figure 4: Stream lines distribution with different roughness r at identical plate distance H = 1.6 and initial Reynolds number Re = 10 on trapezoid substrate. (a) r = 3/8; (b) r = 5/8; (c) r = 6/8; (d) r = 7/8.

Here the definition of basal plate roughness with isosceles trapezoid is r=(H−h)/(W−D) .

Analysis of the Influence of Reynolds Number on the Flow Field of the Thin Film Coating In theory, film coating fluid should maintain in the laminar or nearly laminar flow state, which means that the flow must be carried out as far as possible at a low Reynolds number range [9] . Low Reynolds number means that the viscous force dominates flow process, in other words, that the effect of inertia force can be ignored to some extent [10] . As can be seen from the Figure 4and Figure 5, when the roughness is small such as r < 0.3, no eddy current arises in low Reynolds number range. State of laminar flow is relatively simple; it’s redundant to give details. Here is the only research

Numerical Simulation of the Thin Film Coating Flow in Two-Dimension

11

to produce eddy current flow on the uneven base plate, and this means roughness of the basal plate chosen to discuss is greater than 0.3. Triangle plate and trapezoidal plate flow are similar, so just one of them is chosen to study.

Figure 5: Stream lines distribution with different Reynolds numbers Re at unflatness r = 0.6. (a) Re = 5; (b) Re = 30; (c) Re = 60; (d) Re = 90.

As can be seen from Figure 5, at a low Reynolds range of 1 - 100, when Reynolds number is small, near the basal plate position the volute is small, and flow state is relatively stable. With the increase of Reynolds number, the eddy is more and more obvious to see. But when the Reynolds number increases to a certain extent, volute changes will not be obvious. In other words, compared with the unevenness, the Reynolds number for the eddy current generation and development is not the dominant factor.

Analysis of Thin Film Coating Fluid Velocity Velocity contours of film coating fluid on the triangle basal plate and trapezoid substrate at the same low Reynolds number are respectively

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Modern Techniques of Coating and Drying in Surface Engineering

shown as Figure 6 and Figure 7. Here selected to study is the fluid field on basal plate with high roughness that is greater than 0.5, since there is no need for the research on substrate with a low unevenness (such as r < 0.3) by the reason that the velocity contours of fluid field on substrate at low trapezoid substrate are similar because its state is laminar at low Reynolds number range, while the flow state on the basal plate with larger roughness is turbulent accompanied by segregation phenomenon appearing. As can be seen from the diagram, the fluid velocity close to the bottom is zero, and the overall velocity distribution is symmetrical (in the range of low Reynolds). Whether triangular or trapezoidal substrate, there is an oval center dead zones (backwater region), where the fluid is in static state, and velocity of the surrounding fluid is not equal to zero. Moreover, due to the influence of unevenness, dead zone center arises in a slightly different position, along with the increase of roughness, location has been moved up slightly. There is no center backwater region in the flow field of laminar state, so the center dead zones can be taken as a reference to judge whether the vortex exists or not.

Figure 6: Contours of the velocity on triangle base plate. (a) r = 5/8; (b) r = 1.0.

Figure 7: Contours of the velocity on trapezoid base plate. (a) r = 5/8; (b) r = 7/8.

Numerical Simulation of the Thin Film Coating Flow in Two-Dimension

13

CONCLUSION We proposed a two-dimension model of slip coating under isosceles triangle and isosceles trapezoid substrate, and thin film coating fluid motions under different configuration parameters were simulated. It is pointed out that the key factor determining the turbulence generation and evolution is the parameter of substratum surface nature, with the increase of basal plate roughness, flow separation phenomenon of thin film coating fluid occurs, accompanied by the vortices. Increasing unevenness r will make the eddy current phenomenon more and more obvious, and effect the location of vortex generation. The effects of the change of Reynolds number on turbulent appearance and action area are studied, compared with the unevenness, the Reynolds number for the eddy current generation and development is not the dominant factor. The velocity contours of fluid field on different substrate surfaces are shown, and the impact of substrate geometry on the backwater region is analyzed. The center dead zones exists can be taken as a reference to judge whether the vortex exists.

ACKNOWLEDGEMENTS This work was partly supported by the Chang Jiang Youth Scholars Program of China and grants (51373033 and 11172064) from the National Natural Science Foundation of China to Prof. Xiaohong Qin. As well as “The Fundamental Research Funds for the Central Universities” and “DHU Distinguished Young Professor Program” to her. It also has the support of the Key grant Project of Chinese Ministry of Education (No 113027A). This work has also been supported by “Sailing Project” from Science and Technology Commission of Shanghai Municipality (14YF1405100) to Dr. Hongnan Zhang.

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REFERENCES 1.

Jiang, T.-Q. (2004) Chemical Rheology. East China University of Science and Technology Press, Shanghai, 223-465. 2. Wang, L., Mei, X.-Z. and Ma, J.H. (2006) Application of Thermal Spray Technology on Resistance to Fretting. Nonferrous Metals, 9395. 3. Fu, J.-B. and Zhou, S.-K. (2006) Application of the Thermal Spraying Technology in Aero-Engine Part and Its Service. Failure Analysis and Prevention, 1, 61-64. 4. Li, X.-H., AI, Y.-H. and He, X.-J. (2006) Application of Thermal Spray on Ship Self-Defence Field. Mine Warfare & Ship Self-Defence, 14, 38-41. 5. Edgar, B.G. and Edward, D.C. (1995) Coating and Drying Defects Troubleshooting Operating Problems. A Wiley-Interscience Publication, New York, 4-10, 75-86. 6. Liu, Z.-M., Jin, Y.-M. and Liu, H.-M. (2008) Progress on Formation and Prevention of Defects in Thin Film Coating. Journal of Safety and Environment, 8, 135-139. 7. Taylor, D.J. and Bimie, D.P. (2002) A Case Study in Striation Prevention by Targeted Formulation Adjustment: Aluminum Titanate Sol-Gel Coatings. Chemistry of Materials, 14, 1488-1492. https://doi. org/10.1021/cm010192c 8. Steven, J.W. and Chen, K.-P. (1999) Large Growth Rate Instabilities in Three-Layer Flow down an Incline in the Limit of Zero Reynolds Number. Physics of Fluids, 11, 3270-3282. https://doi. org/10.1063/1.870187 9. Scholle, M., Wierschem, A. and Aksel, N. (2004) Creeping Films with Vortices over Strongly Undulated Bottoms. Acta Mechanica, 168, 167193.https://doi.org/10.1007/s00707-004-0083-4 10. Wierschem, A. and Aksel, N. (2004) Creeping Films with Vortices over Strongly Undulated Substrates. Physics of Fluids, 12, 58-73.

Chapter 2

Corrosion Study of Powder-Coated Galvanised Steel

Manish Kumar Bhadu, Akshya Kumar Guin, Veena Singh, and Shyam K. Choudhary R&D and Scientific Services, TATA Steel Ltd., Jamshedpur 831001, India

ABSTRACT In general, steel is protected from corrosive environments by conversion coatings, that is, phosphating, chromating, and so forth, and then followed by different layers of paints. Nowadays, strict pollution laws and regulations are creating significant challenges for coating experts to develop an environmentally friendly product. Powder coatings have demonstrated their ability as alternative to traditional solvent-borne coatings. In the present work, polyester-based two coating systems have been investigated and their performances have been evaluated for surface topographical properties by Scanning electron microscope (SEM), and energy dispersive spectrometry

Citation: Manish Kumar Bhadu, Akshya Kumar Guin, Veena Singh, and Shyam K. Choudhary, “Corrosion Study of Powder-Coated Galvanised Steel,” ISRN Corrosion, vol. 2013, Article ID 464710, 9 pages, 2013. https://doi.org/10.1155/2013/464710 Copyright © 2013 Manish Kumar Bhadu et al. This is an open access article distributed under the Creative Commons Attribution License, which permits unrestricted use, distribution, and reproduction in any medium, provided the original work is properly cited.

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Modern Techniques of Coating and Drying in Surface Engineering

(EDX), accelerated corrosion resistance by salt spray test, and impedance property by electrochemical impedance spectroscopy. Coating adhesion with the steel and hardness were evaluated by bond strength, cross cut adhesion, and pencil hardness. This paper explains the results and performance of the coatings by the above two systems.

INTRODUCTION Galvanised steel is widely used in construction, automobile, and white goods sector. Zinc coating is the most effective and economic means to protect the steel substrate exposed to atmospheric corroding environments. It protects the steel substrate by acting as a barrier against the corrosive environment and by sacrificially corroding themselves to provide cathodic protection. Protective ability of galvanised (GI) steel may be enhanced by employing thicker zinc coatings or by painting the metallic substrate [1]. Paints improve the surface life of underlying zinc coating acting as a barrier against zinc reaction with environmental agents. Cracks, crater, and pin holes occurring in the paint are sealed by corroded zinc products. Moreover, corroded zinc products occupy a 20–25% more volume than zinc, while iron oxides (corrosion product of steel) occupy a volume several times larger than the steel; thus, expansive forces are reduced at the zinc-paint interface compared to those at the steel-paint one [2]. The main practical problem concerning painting of zinc-coated surface lies in achieving good bond strength, that is, good adhesion of polymer with GI sheet. Often adhesion looks satisfactory immediately after painting, but it prematurely degrades after water, oxygen, and other corrosive ingredients diffuse through the polymeric coating. Several pretreatment processes were reported to improve coating adhesion. The main function of pretreatment for GI steel surface is to form a very stable passive film which will enhance the adhesion with subsequent polymeric film [3]. There are a number of pretreatment processes like phosphating, chromating, and coating of rust preventing compounds by chemical and physical vapor deposition and diffusion coating, and so forth. Some of these pretreatment processes do not contain a film forming material that is, nonpriming, and hence, they should be compatible with the rest of the painting system [4]. Primers contain a film forming material and are expected to act as the anchorage of the paint system [5]. Some of these coatings are toxic and pose concerns to the environment. Powder thermoset coatings are solvent-free and unlike the conventional liquid coatings have zero volatile organic content (VOC). Thus, they could offer the coating formulators a

Corrosion Study of Powder-Coated Galvanised Steel

17

robust and a highly promising approach to produce eco-friendly coatings. Moreover, not only powder coatings are easier to apply than solvent-based coatings, but, they also provide a thicker and more uniform coating. If we take into account the economical advantages of powder coatings, like cost savings in water disposal, high yield, and cheap maintenance, the total operating cost of a powder application plant is lower than that of traditional solvent-based liquid paints. Furthermore, a wide variety of finishes such as structured, wrinkled, metallic, and antique finishes are available with powder coatings [6–9]. Although it is possible to achieve high-gloss and smooth coatings with excellent adhesion, flexibility, and hardness by epoxy powder coating, they however exhibit poor tolerance to heat and light resulting in a pronounced tendency to fade [10]. In order to overcome this shortcoming, polyester/ epoxy blends-based coating technology has been developed which shows excellent film smoothness, appropriate mechanical properties, and adherence characteristics [11]. Interestingly, blend coatings show a broader degradation temperature interval, and they volatilize to a substantially less extent compared to pure epoxy and polyester powder coating. Epoxypolyester powder coating is a hybrid of epoxy and polyester powder coating. These hybrids have properties similar to those of epoxy powders; however, their additional advantage is that they have improved resistance to fading and improved weather resistance [11]. Hybrid powders are now regarded as the main backbone of the powder coating industry. The objective of the present work is to find out a suitable powder coating system for GI surface. A comparative study was made in between polyester and epoxy-polyester powder coating process. Adhesion on GI steel was assessed by suitable standardized adhesion tests. Anticorrosive behaviour was evaluated by salt spray and electrochemical impedance spectroscopy (EIS) method.

EXPERIMENTAL DETAILS Zero-spangle GI (120 g/m2) sheet was used as the base substrate. These sheets were degreased to remove oil and grease followed by phosphating in tricationic phosphate base solution to obtain a thin coating of dry film thickness (DFT) ~4 micron. The polyester and epoxy-polyester powder coating of DFT 60 micron was applied by electrostatic spraying method on phosphate GI sheet referred to as system-1 and system-2, respectively.

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Modern Techniques of Coating and Drying in Surface Engineering

Coated panels were allowed to stand for 7 days at room temperature for curing before any testing. The microstructure and surface morphology of the coated samples were observed by a scanning electron microscope (SEM) equipped with energy dispersive spectrometry (EDX). The energy used for analysis was 15 KeV. Corrosion resistance of the samples was evaluated by electrochemical impedance spectroscopy (EIS). The EIS measurements were carried out by using the VersaSTAT MC. A typical three-electrode system was employed in these tests. The samples acted as the working electrode (1 cm2 of the exposed area), saturated calomel electrode (SCE) as the reference, and graphite as counter electrode. 3.5% NaCl solution was used as an electrolyte in all the measurements. The EIS measurement was carried out in the frequency range of 100 KHz to 0.01 Hz, and the applied voltage was 5 mV. Salt spray tests were carried out by exposing the scribed samples (6’’×4’’) in a salt spray chamber as per the ASTM B-117 test method. The panels were checked at a regular interval of time, and results were noted down in terms of blisters, creep, and red rust. The adhesion, hardness, and bond strength measurements were performed on the coated steel sample as per the ASTM D3359, ASTM D 3363, and ASTM D 4541, respectively.

RESULTS AND DISCUSSION Morphology The schematic diagram of the different coating layers on the steel substrate is shown in Figure 1. The cross-sectional SEM photographs of phosphate steel with polyester powder coated and epoxy-polyester powder coated are shown in Figures 2 and 3, respectively. It can be seen that coating formed by epoxy-polyester appears to be uniform, without any surface defects and cracks. From the SEM photograph, it is clearly visible that the epoxypolyester powder coating is highly dense, and it strongly adheres to surface, whereas some visible cracks appear on polyester-coated sheet and are not so strongly adherent to surface compared to the epoxy-polyester-coated surface.

Corrosion Study of Powder-Coated Galvanised Steel

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ISRN Corrosion

Figure 1: Schematic diagram of different coating layers with thickness.

3

Fe

Fe

Fe Zn

C

0

1

Zn

2

3

4

5

6

7

8

9

Zn

10

(keV) Full-scale 5718 cts cursor: −0.018 (353 cts)

Electron image 1

90 𝜇𝜇m

Point 1

(a)

(b)

C

Zn

S O

O Ti Fe

Fe

S

C

0

Ba

Fe

1

2

Ti Ti

3

Zn

Fe

Zn

Ba

Si

Ba

Zn

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5 6 7 8 9 10 (keV) Full-scale 5718 cts cursor: −0.018 (297 cts) Point 3 (c)

1

2

3

4

Ba BaBa Fe

5 6 (keV)

Fe

7

Zn

8

Zn

9

10

Full-scale 1613 cts cursor: 0.048 (306 cts) Point 7

(d)

Figure 2: (a) Cross-sectional SEM image of polyester powder-coated sheet (system-1). (b), (c), and (d) show EDX analysis of three respective points 1, 3, and 7, shown in (a). Figure 2: (a) Cross-sectional SEM image of polyester powder-coated sheet

(system-1). (b), (c), and (d) show EDX analysis of three respective points 1, 3, and 7, shown in (a). the case of system-2 is from the glue itself. such as defects, adhesion, and barrier properties,

The failure in But polyester coating system shows failure from interface and needs only 4 MPa force for any type of coating delamination. This improvement in system-2 is due to the formation of strong and crack-free bonding of epoxy-polyester hybrid with steel substrate, whereas polyester coating weakly adheres with the presence of internal crack, as noticed from SEM photograph and hence needs minimum force to delaminate from the surface as compared to epoxy-polyester coating [12–

as well as to determine the onset and progression of corrosion process on metal substrate underneath the organic coating. For the EIS measurements, elements were selected through a model equivalent circuit (see Figure 4) to represents the systems under study. The systems undergo a charge transfer control according to this circuit, where 𝑅𝑅s represents solution resistance, 𝑅𝑅p coating resistance, 𝐶𝐶c coating capacitance, 𝑅𝑅ct charge transfer resistance, and 𝐶𝐶dl double layer capacitance.

ISRN Corrosion

20

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Modern Techniques of Coating and Drying in Surface Engineering

Zn O Mn Fe P

S Ba

Zn

Ba Ni

1

Mn Ba Fe Ba Mn Ba

Ba

2

3

4

5

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Fe Ni

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Zn

Ni

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9

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(keV) Full-scale 785 cts cursor: 0.028 (359 cts)

Electron image 1

90 𝜇𝜇m

Point 6

(a)

(b)

Fe

Zn

Fe

Zn

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Fe

Fe

O Fe

Fe

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Zn

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5 6 7 8 9 10 (keV) Full-scale 4830 cts cursor: 0.028 (359 cts) Point 7 (c)

1

2

3

4

5 6 7 8 9 10 (keV) Full-scale 3118 cts cursor: 0.093 (196 cts) Point 9 (d)

Figure 3: (a) Cross sectional SEM image of epoxy-polyester powder-coated sheet (system-2). (b), (c) and, (d) show EDX analysis of three (a) inCross sectional SEM image of epoxy-polyester powder-coated respective pointsFigure 6, 7, and 9,3: shown (a).

sheet (system-2). (b), (c) and, (d) show EDX analysis of three respective points 𝐶𝐶c 6, 7, and 9, shown in (a). 𝑅𝑅s

𝐶𝐶dl

𝑅𝑅p EDX spectra of polyester powder-coated and epoxy-polyester powdercoated samples are also shown in Figures 2 and 3. The peaks of different 𝑅𝑅ct pigments materials like zinc, silica, phosphorus, and oxygen are predominant Figure 4: Electrical equivalent circuit for the polyester and polyester-epoxy coating on the galvanised sheet. in both cases.

Bond Strength

and steel substrate create an impermeable surface for water and other corrosive ingredients on the coated surface [18–20]. Two major pieces of information

epoxy-polyester powder-coated sample (system-2) is found to have less white rust even after 1608 h of exposure in spray chamber 117). Whereas coating systemaresalt obtained from(ASTM the Bbond strength 1 containing polyester powder coating fails after 1200 h of The Property first is bytheSaltpull-off strength, that is, bond strength of coating on There are a exposure in salt spray chamber (ASTM B117). 3.4. Corrosion test. Resistance Sprat Test. Pernumber of blisters and split around the edges of formance of steel coated sheet, h exposure salt substrate, andafter the1608 second one isin about the point whereinthe occurred inthe scribe area in the case of system-1. spray chamber is shown in Figure 11. It can be seen that

the paint system. The split could be an adhesive break, a cohesive break, a

Corrosion Study of Powder-Coated Galvanised Steel

21

combination of both, or a failure of the glue. From Table 1, cohesive failure is observed in polyester powder-coated sample, whereas adhesive failure is observed in epoxy-polyester system. Epoxy-polyester system comparatively adheres strongly with phosphated steel substrate and needs more than 5 MPa force for any type of delamination. The failure in the case of system-2 is from the glue itself. But polyester coating system shows failure from interface and needs only 4 MPa force for any type of coating delamination. This improvement in system-2 is due to the formation of strong and crack-free bonding of epoxy-polyester hybrid with steel substrate, whereas polyester coating weakly adheres with the presence of internal crack, as noticed from SEM photograph and hence needs minimum force to delaminate from the surface as compared to epoxypolyester coating [12–14]. Table 1: Bond strength and nature of failure of galvanised coated substrate System

Nature of failure

1

Bond strength (in MPa) 4

2

5

Failure from top coat and adhesive

Photographs

Failure from top coat and adhesive

Corrosion Resistance Property by Electrochemical Impedance Spectroscopy The electrochemical properties of polyester and epoxy-polyester on phosphate steel substrate were examined by EIS measurements. The main role of EIS in the characterization of an organic coating is to provide information about the properties of the protective system, such as defects, adhesion, and barrier properties, as well as to determine the onset and progression of corrosion process on metal substrate underneath the organic coating. For the EIS measurements, elements were selected through a model

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Modern Techniques of Coating and Drying in Surface Engineering

equivalent circuit (see Figure 4) to represents the systems under study. The systems undergo a charge transfer control according to this circuit, where Rs represents solution resistance, Rp coating resistance, Cc coating capacitance, Rct charge transfer resistance, and Cdl double layer capacitance.

Figure 4: Electrical equivalent circuit for the polyester and polyester-epoxy coating on the galvanised sheet.

The Bode representations of the impedance data have been analyzed with VersaSTAT MC and ZSimpWin software of Princeton applied research. Single slope in the midfrequency range shows the existence of a single time constant, and the impedance data have been analyzed using the equivalent circuit. The impedance behavior of coating system-1 and coating system-2 on steel substrate after initial study, 144 h, and 264 h of immersion in 3.5% NaCl solution is shown Figures 5(a)–5(c) with overlay. The individual impedance behaviours with initial time, 24 h, 96 h, 144 h, and 264 h are shown in Figures 6, 7, 8, 9, and 10, respectively. The coating resistance and capacitance values derived from these figures are given in Table 2. It is clear from Figures 6, 7, 8, 9, and 10 and Table 2 that in the beginning of experiment, the coating resistance of polyester- and epoxy-polyester-coated samples is in the same range. Impedance of coating system-1 containing polyester powder coating shows 8.835×106Ω/cm2 against the 1.166×106 Ω/ cm2 resistance of coating system-2. Table 2: Impedance parameters of wash primer and epoxy primer and top coated GI in 3.5% NaCl solution Time, h Initial 24 h

System-1 resistance (Ω/cm2) 8.835 × 106 7.203 × 104

System-2 resistance (Ω/cm2) 1.166 × 106 1.446 × 105

Corrosion Study of Powder-Coated Galvanised Steel 96 h 144 h 264 h

7367 7166 5504

23

1.185 × 105 1.633 × 104 1.576 × 104

Figure 5: (a), (b), and (c) show EIS overlap diagram of samples immersed for time intervals, beginning, 144 h and 264 h respectively.

Figure 6: Bode plot of (a) system-1 and (b) system-2 at initial time.

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Modern Techniques of Coating and Drying in Surface Engineering

Figure 7: Bode plot of (a) system-1 and (b) system-2 at 24 h time.

Figure 8: Bode plot of (a) system-1 (b) system-2 at 96 h time.

Figure 9: Bode plot of (a) system-1 and (b) system-2 at 144 h time.

Figure 10: Bode plot of (a) system-1 and (b) system-2 at 264 h time.

Corrosion Study of Powder-Coated Galvanised Steel

25

The coating resistance value of system-1 decreases from 8.835×106 Ω/cm2 to 5.504 × 103 Ω/cm2 after 264 h of immersion, whereas coating resistance value of system-2 decreases from 1.166 × 106 Ω/cm2 to 1.633 × 104 Ω/cm2 in 144 h of immersion and to 1.576 × 104 Ω/cm2 in 264 h of immersion. Generally, the high impedance value of polyester-coated sample shows a fast reduction in the first 24 h of immersion, due to the development of conductive pathways inside the film. Comparatively, a slow decrease in the impedance value was observed for the epoxy-polyester powder-coated steel sheet, followed by a small recovery after 24 h of immersion [15, 16]. Penetration of water and movement of ionic species through the coating layer may be responsible for the observed decrease in coating resistance value [17]. Other reasons could be a weaker ionic resistance and a lower-cross linking density. The dielectric constants of organic coating and water are about 6 and 80, respectively, at ambient temperature. Therefore, permeation of a small amount of water through the coating can contribute to a relatively large change in the pore resistance. It is known that diffusion of electrolyte, water, and ions through the epoxy-polyester powder-coated sample is much lower than the polyester powder-coated sample. The higher corrosion protection by epoxy-polyester powder coating on steel surface is due to the higher cross-linked density of epoxy-polyester network comparing with polyester system; in addition to high cross-linking density, the additional free hydroxyl group, that is, –OH of epoxy-polyester coating system (as compared to only polyester system), forms a strong bond with phosphate and steel surface. The high cross-linking density and strong adhesion of epoxy group with phosphate and steel substrate create an impermeable surface for water and other corrosive ingredients on the coated surface [18–20].

Corrosion Resistance Property by Salt Sprat Test Performance of steel coated sheet, after 1608 h exposure in salt spray chamber is shown in Figure 11. It can be seen that epoxy-polyester powdercoated sample (system-2) is found to have less white rust even after 1608 h of exposure in salt spray chamber (ASTM B 117). Whereas coating system1 containing polyester powder coating fails after 1200 h of exposure in salt spray chamber (ASTM B117). There are a number of blisters in and around the edges of the scribe area in the case of system-1.

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Modern Techniques of Coating and Drying in Surface Engineering

Figure 11: Salt spray photographs of coated substrate, (a) polyester-coated sample after 1200 h of salt spray test, (b) epoxy-polyester-coated sample after 1608 h of the salt spray test.

Adhesion and Hardness Cross hatch and pencil hardness of system-1 and system-2 were performed according to the ASTM D3359 and ASTM D3363-05. The results of these tests are shown in Table 3, which compares the adhesion and hardness properties of the two systems. Table 3: Cross-hatch adhesion and pencil hardness results of system-1 and system-2 System 1 2

Cross-hatch 5B 5B

Percentage of area removed 0 0

Pencil hardness (at 45°) 7H 9H

The result of the x-cut adhesion test was satisfactory for both the systems providing 5B with no observed flaking in the cross-cut area. Therefore, no adhesion losses appeared in any interfaces. The adhesion of primer to the steel surface and the adhesion of the top coats to the primer, both were satisfactory.

Corrosion Study of Powder-Coated Galvanised Steel

27

CONCLUSION The powder coatings protect galvanised steel substrates by the introduction of a barrier layer with relatively high ohmic resistance between the metallic substrate and the corrosive environment. Also, it is indicated that the crosslink density appears only to affect ionic conductivity of the film, and due to the higher cross-linked density of epoxy-polyester network comparing with polyester system, the permeability of water and ions through the coating film becomes less, and it leads to a more impervious film with a more resistant structure to corrosion. It is indicated that the polyester film is more porous due to the less cross-link density of cured polymer, compared with epoxy-polyester powder-coated samples. SEM micrograph and bond strength suggest that epoxy-polyester coating is more firmly adherent to the phosphate substrate than polyester coating due to more hydroxyl groups in epoxy-polyester bonding with phosphate substrate. Protective properties of fully cured organic coatings on metallic substrate may be attributed to a barrier and/or an active inhibition mechanism.

ACKNOWLEDGMENT The authors would like to thank the staff of the corrosion lab for helping them during experiments.

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REFERENCES 1.

A. K. Guin, S. Nayak, T. K. Rout, N. Bandyopadhyay, and D. K. Sengupta, “Corrosion resistance nano-hybrid sol-gel coating on steel sheet,” ISIJ International, vol. 51, no. 3, pp. 435–440, 2011. 2. A. S. Khanna, Introduction to High Temperature Oxidation and Corrosion, ASM International, 1997. 3. B. V. Jegdić, J. B. Bajat, J. P. Popić, and V. B. Mišković-Stanković, “Corrosion stability of polyester coatings on steel pretreated with different iron-phosphate coatings,” Progress in Organic Coatings, vol. 70, no. 2-3, pp. 127–133, 2011. 4. A. M. P. Simões, R. O. Carbonari, A. R. Di Sarli, B. del Amo, and R. Romagnoli, “An environmentally acceptable primer for galvanized steel: formulation and evaluation by SVET,” Corrosion Science, vol. 53, no. 1, pp. 464–472, 2011. 5. R. Mafi, S. M. Mirabedini, M. M. Attar, and S. Moradian, “Cure characterization of epoxy and polyester clear powder coatings using Differential Scanning Calorimetry (DSC) and Dynamic Mechanical Thermal Analysis (DMTA),” Progress in Organic Coatings, vol. 54, no. 3, pp. 164–169, 2005. 6. D. Maetens, “Weathering degradation mechanism in polyester powder coatings,” Progress in Organic Coatings, vol. 58, no. 2-3, pp. 172–179, 2007. 7. J. B. Bajat, J. P. Popić, and V. B. Mišković-Stanković, “The influence of aluminium surface pretreatment on the corrosion stability and adhesion of powder polyester coating,” Progress in Organic Coatings, vol. 69, no. 4, pp. 316–321, 2010. 8. M. A. J. Batista, R. P. Moraes, J. C. S. Barbosa, P. C. Oliveira, and A. M. Santos, “Effect of the polyester chemical structure on the stability of polyester-melamine coatings when exposed to accelerated weathering,” Progress in Organic Coatings, vol. 71, no. 3, pp. 265–273, 2011. 9. F. L. Duivenvoorde, J. Laven, and R. Van der Linde, “Diblock copolymer dispersants in polyester powder coatings,” Progress in Organic Coatings, vol. 45, no. 2-3, pp. 127–137, 2002. 10. S. Radhakrishnan, N. Sonawane, and C. R. Siju, “Epoxy powder coatings containing polyaniline for enhanced corrosion protection,” Progress in Organic Coatings, vol. 64, no. 4, pp. 383–386, 2009.

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11. V. C. Malshe and G. Waghoo, “Weathering study of epoxy paints,” Progress in Organic Coatings, vol. 51, no. 4, pp. 267–272, 2004. 12. R. Van der Linde, E. G. Belder, and D. Y. Perera, “Effect of physical aging and thermal stress on the behavior of polyester/TGIC powder coatings,” Progress in Organic Coatings, vol. 40, no. 1–4, pp. 215–224, 2000. 13. Technical Data Sheet of Ridoline 1352 BA, of M/s Henkel Chembond India Ltd. 14. Technical Data Sheet of Granodine, of M/s Henkel Chembond India Ltd. 15. M. Özcan, I. Dehri, and M. Erbil, “EIS study of the effect of high levels of SO2 on the corrosion of polyester-coated galvanised steel at different relative humidities,” Progress in Organic Coatings, vol. 44, no. 4, pp. 279–285, 2002. 16. R. Naderi, M. M. Attar, and M. H. Moayed, “EIS examination of mill scale on mild steel with polyester-epoxy powder coating,” Progress in Organic Coatings, vol. 50, no. 3, pp. 162–165, 2004. 17. R. Mafi, S. M. Mirabedini, R. Naderi, and M. M. Attar, “Effect of curing characterization on the corrosion performance of polyester and polyester/epoxy powder coatings,” Corrosion Science, vol. 50, no. 12, pp. 3280–3286, 2008. 18. G. Wuzella, A. Kandelbauer, A. R. Mahendran, and A. Teischinger, “Thermochemical and isoconversional kinetic analysis of a polyesterepoxy hybrid powder coating resin for wood based panel finishing,” Progress in Organic Coatings, vol. 70, no. 4, pp. 186–191, 2011. 19. M. Barletta and A. Gisario, “An application of neural network solutions to laser assisted paint stripping process of hybrid epoxy-polyester coatings on aluminum substrates,” Surface and Coatings Technology, vol. 200, no. 24, pp. 6678–6689, 2006. 20. R. Naderi, M. M. Attar, and M. H. Moayed, “EIS examination of mill scale on mild steel with polyester-epoxy powder coating,” Progress in Organic Coatings, vol. 50, no. 3, pp. 162–165, 2004.

Chapter 3

Toward a Nearly Defect-free Coating via High-energy Plasma Sparks

Mosab Kaseem, Hae Woong Yang & Young Gun Ko Materials Electrochemistry Laboratory, School of Materials Science and Engineering, Yeungnam University, Gyeongsan, 38541, Republic of Korea

ABSTRACT A nearly defect-free metal-oxide-based coating structure was made on Al-Mg-Si alloy by plasma electrolytic oxidation at high current density accompanying high-energy plasma sparks. The present coatings were performed at two different current densities of 50 and 125 mA/cm2 in the

Citation: Mosab Kaseem, Hae Woong Yang & Young Gun Ko “Toward a nearly defect-free coating via high-energy plasma sparks” Scientific Reports volume 7, Article number: 2378 (2017). https://doi.org/10.1038/s41598-017-02702-3 Copyright © The Author(s) 2017. This article is licensed under a Creative Commons Attribution 4.0 International License, which permits use, sharing, adaptation, distribution and reproduction in any medium or format, as long as you give appropriate credit to the original author(s) and the source, provide a link to the Creative Commons license, and indicate if changes were made. The images or other third party material in this article are included in the article’s Creative Commons license, unless indicated otherwise in a credit line to the material. If material is not included in the article’s Creative Commons license and your intended use is not per-mitted by statutory regulation or exceeds the permitted use, you will need to obtain permission directly from the copyright holder. To view a copy of this license, visit http://creativecommons.org/licenses/by/4.0/.

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alkaline-phosphate-based electrolytes with different concentrations of sodium hexafluoroaluminate (Na3AlF6). The addition of (Na3AlF6) to the electrolyte used in this study would result in a decrease in the size of the micropore, and a reasonably defect-free coating structure was achieved in the sample treated at high current density of 125 mA/cm2. This was attributed mainly to the hydrolysis of AlF6 3− triggered by intense plasma sparks, which resulted in a uniform distribution of fluorine throughout the coating. Accordingly, the corrosion performance of the coating formed in the electrolyte containing 1.5 g/L Na3AlF6 at 125 mA/cm2 was improved significantly as confirmed by electrochemical impedance analysis. In addition, the formation mechanism of the nearly defect-free coating in the presence of Na3AlF6 was discussed.

INTRODUCTION In recent years, a plasma electrolytic oxidation (PEO) has emerged as a novel method for surface treatment capable of producing metal-oxidebased coatings with desirable properties on the surface of valve metals and alloys1,2,3,4,5,6. When a high voltage above the breakdown voltage would be applied, a protective coating formed on the surface of the metal substrate by plasma-assisted electrochemical reactions7, 8. Typically, the structure of PEO coating comprised two distinct layers: an outer layer with a number of the micropores, cracks, and structural defects, and an inner layer with a relatively compact structure. The micropores and other defects were distributed inhomogeneously in both layers. They were more evident in the outer layer. This would cause the infiltration of corrosive medium into the inner layer of the coating and, thereby, arrive at the substrate, which accelerated the corrosion rate due to a local change in pH9. To alleviate the present problem, considerable attention has been paid to improving the fabrication of defect-free coatings by PEO through the development of novel electrolytic systems together with the optimization of the electrical parameters, such as frequency and current density8. Among various approaches, the incorporations of inorganic particles, such as ZrO2, TiO2, CeO2, and clay, was considered one of the useful strategies, Kaseem et al.10 suggested an effective way to block the micropores by incorporating ZrO2 and MoO2 particles into the coating, which improved significantly the corrosion performance of 7075 Al alloy. According to recent results by Rapheal et al.11, however, the micropores and cracks were still clearly visible in the coating, which might provide the short paths for

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the corrosive medium although the addition of clay helped increase the compactness of the coating formed on an AM50 Mg alloy by PEO at a low current density of 30 mA/cm2. In contrast, Lu et al.12 demonstrated that, in coating formed by PEO on AM50 Mg alloy, the micropores would be filled fairly by adding Si3N4particles to the KOH-Na3PO4-containing electrolyte. The distribution of the microdefects found in cross section of the coating was still in argument. Thus, the improvement in corrosion performance was case-sensitive. On the other hand, the incorporation of F− ions into the coatings has been reported as an effective strategy to improve the corrosion performance13,14,15,16,17,18. Per potentiodynamic polarization results reported by Kazanski et al.13, the coating formed on AZ91 Mg alloy by PEO under alternating current in an alkaline silicate-based electrolyte containing KF exhibited better corrosion performance as compared to the counterpart coating formed without KF. This result suggested that a significant drop in the porosity was achieved by the incorporation of F− ions into the coating. Indeed, Duan et al.18 suggested that KF was able to enhance the corrosion resistance of the inner layer. It was, therefore, concluded that F− ions would improve the corrosion performance due to the formations of F-compounds that would increase the thickness and compactness of the inner layer. In case of Mg alloy, MgF2 would precipitate readily on the surface of the anode, preventing excessive dissolution of Mg element from the substrate8, 14. However, Yerokhin et al.19 reported that the addition of F− ions to the electrolyte showed a different tendency of Al-oxide coating. Sodium hexafluoroaluminate (Na3AlF6) was prone to be hydrolyzed in aqueous solution, resulting in the formations of aluminum hydroxide (Al(OH)3) and F-containing compounds. Hence, the observations in the previous studies raised the question of whether the use of the complex fluorine-containing salts would be desirable for fabricating a nearly defectfree coating via PEO because the hydrolysis of Na3AlF6 might lead to the structural modification of the coating. Up to the present study, the fabrication of a nearly defect-free coating on Al-based alloy under high-energy plasma condition using the fluorine-containing salt as an additive in the electrolyte and the electrochemical contribution of the present coating have rarely been understood. Therefore, the aim of this work is to fabricate the nearly defectfree compact coating on Al-Mg-Si alloy by PEO to improve corrosion performance by adding Na3AlF6 to the alkaline phosphate-based electrolyte. On the basis of the results, the mechanism on the formation of the coatings

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in the presence of Na3AlF6 will be discussed in relation to the structural characteristics controlled with respect to current density as well as Na3AlF6 concentration.

RESULTS AND DISCUSSION Structural Characteristics of Coating Figure 1 shows the changes in surface morphology of the PEO coatings as a function of Na3AlF6 concentration and current density. In all coatings, the surface was dominated by numerous micropores, which are characteristic of PEO coatings. To elucidate the effect of Na3AlF6concentration and current density on the surface morphology of the coatings formed by PEO, the porosity and pore size values were measured, and they are listed in Table 1. Irrespective of current density, it was observed that the average size of the micropores and cracks, as well as the degree of porosity, was lower for the coatings formed with 1.5 g/L Na3AlF6. This suggested that the addition of 1.5 g/L Na3AlF6 to the phosphate-based electrolyte might decrease the intensity of plasma discharges, resulting in the generation of a compact microstructure with less structural defects (Fig. 1b and e). Indeed, when adding 3 g/L Na3AlF6 to the electrolyte, the porosity and micropore size increased as compared to the sample obtained in an electrolyte containing 1.5 g/L Na3AlF6. In addition, from the high-magnification SEM images shown in the insets of Fig. 1, two important observations were made regarding surface morphology. Firstly, as current density increased, the micropore size tended to increase, whereas the percentage of porosity tended to decrease. Secondly, the coating obtained with no Na3AlF6 at 125 mA/cm2 exhibited severe cracking as compared to the coatings formed in electrolytes containing Na3AlF6, which is consistent with earlier reports11, 20, 21. For the coatings formed in the electrolyte with no Na3AlF6, larger micropores and cracks were formed at the higher current density due to the increased plasma discharge intensity after quenching by the surrounding electrolyte. However, interestingly, for the coatings formed in electrolyte containing Na3AlF6, the number of cracks tended to decrease as current density increased, suggesting that the addition of Na3AlF6 together with a higher current density could promote the incorporation of F-containing compounds, thus decreasing the size of micropores and cracks (Fig. 1e and f).

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Figure 1: SEM images showing the surface morphology of the coatings formed on an Al-Mg-Si alloy by PEO at 50 mA/cm2 (a–c) and 125 mA/cm2 (d–f) in an electrolyte containing no Na3AlF6 (a and d), 1.5 g/L Na3AlF6 (b and e), and 3 g/L Na3AlF6 (c and f). Insets are the high-magnification images of the surface morphology of the PEO coatings. Table 1: Pore size and porosity of the PEO coatings formed on an Al-Mg-Si alloy by PEO for 8 min using different concentrations of Na3AlF6 at different current densities

Table 2 lists the EDS results corresponding to the surface of the coatings obtained in different electrolytes at different current densities. Irrespective of current density, EDS results showed that the coatings formed in the electrolyte with no Na3AlF6 were mainly composed of Al, O, and P, while as expected, F was only detected in the coatings formed in electrolytes containing Na3AlF6. This result indicated that AlF6 3− ions effectively contributed to form the coating layers during the plasma-assisted electrochemical reaction. It is worth noting that the amount of P was observed to decrease in the coatings formed in electrolytes containing Na3AlF6, which was attributed to

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the fact that AlF6 the surface22.

3−

ions prevented the adsorption of phosphate groups on

Table 2: EDS results of the PEO coatings formed on an Al-Mg-Si alloy by PEO for 8 min using different concentrations of Na3AlF6 at different current densities

The cross-sectional images of the PEO coatings obtained in different electrolytes and at different current densities are shown in Fig. 2. It was observed that Na3AlF6 had no effect on the thickness of the coating. By contrast, the coatings obtained at a current density of 50 mA/cm2exhibited a thickness of about 10 µm, while the coatings obtained at 125 mA/cm2 were thicker (with a thickness of ~15 µm). These results were reasonably consistent with those obtained for Ti alloy samples coated by PEO in electrolytes containing different concentrations of calcium hypophosphite23. In addition, it was clear that the coatings obtained with Na3AlF6 exhibited a more compact structure as compared to their counterparts obtained in the electrolyte with no Na3AlF6, indicating that AlF6 3− ions decreased the porosity and enhanced the density of the PEO coatings. Since the inner layer of the coatings formed in electrolyte containing Na3AlF6 was thicker and more compact, it was expected that such coatings would exhibit a higher resistance against corrosion. The elemental distribution along the cross-section of the coatings formed at 125 mA/cm2 in electrolytes either with no Na3AlF6 or containing 3 g/L Na3AlF6 was investigated by EDS line scan (Fig. 3). The EDS analysis of the PEO coating obtained with 3 g/L Na3AlF6 clearly showed that F was almost uniformly distributed along the cross-section of the coating, while the distribution of P was less uniform as compared to the counterpart coating obtained without Na3AlF6.

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Figure 2: Cross-sectional images of the coatings formed on an Al-Mg-Si alloy by PEO at 50 mA/cm2 (a–c) and 125 mA/cm2 (d–f) in an electrolyte containing no Na3AlF6 (a and d), 1.5 g/L Na3AlF6 (band e), and 3 g/L Na3AlF6 (c and f).

Figure 3: EDS line scan of the cross-section of the coatings formed on an AlMg-Si alloy at 125 mA/cm2in electrolyte containing no Na3AlF6 (a–c) and 3 g/L Na3AlF6 (d–g).

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Figure 4 shows the XRD patterns of the PEO coatings obtained in different electrolytes at different current densities. The XRD peaks corresponding to Al, γ-Al2O3, α-Al2O3, and AlPO4 were detected in all coatings. The identification of metallic Al in the pattern of the PEO coatings was ascribed to porosity, which caused X-ray penetration through the Al-Mg-Si alloy substrate. During PEO, Al3+ ions were released from the substrate and reacted with O2− ions in the electrolyte to produce γ-Al2O3 and α-Al2O3 on the surface of the Al-Mg-Si alloy at rapid and slow cooling rates, respectively24, 25. On the other hand, AlPO4could be formed by the reaction between PO4 3− ions, generated through ionization of Na3PO4 in the electrolyte, and Al+3 ions (Equation 1). It is worth noting that although the EDS analysis showed a significant amount of F in the coatings formed in electrolytes containing Na3AlF6, no peaks attributed to F-containing phases were detected in the XRD patterns, suggesting that F-containing compounds were incorporated into the coatings as the amorphous phase, rather than the crystalline. In addition, it was observed that in the XRD pattern of the coatings formed in electrolyte containing Na3AlF6, the intensity of all characteristic peaks of γ-Al2O3 was higher. This suggested that the addition of Na3AlF6 to the phosphate-based electrolyte affected the phase composition of the coatings. Furthermore, for all coatings, the amount of α-Al2O3 tended to be lower as compared to that of γ-Al2O3, which was attributed to the fact that the rapid solidification of molten Al2O3 did not allow the complete transformation from metastable γ-Al2O3 into stable α-Al2O3. Therefore, the amount of the α-Al2O3 phase obtained was lower under these experimental conditions26. Finally, according to the XRD and EDS results, the possible reactions during PEO in the different electrolytes are the following: (1) (2) (3) (4)

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Figure 4: XRD patterns of the coatings formed on al Al-Mg-Si alloy by PEO at (a) 50 mA/cm2 and (b) 125 mA/cm2 using electrolytes containing different concentrations of Na3AlF6.

Electrochemical Behavior of Coating The corrosion resistance of the PEO coatings formed in various electrolytes at two different current densities was evaluated by potentiodynamic polarization tests, and curves are shown in Fig. 5. Corrosion parameters such as the corrosion potential (E corr ) and corrosion current density (i corr ) were extrapolated from the polarization curves, and the results are listed in Table 3. These results indicated that all coatings obtained in electrolytes containing Na3AlF6exhibited a higher resistance against corrosion as compared to the coatings formed in the electrolyte with no Na3AlF6. In particular, the coating formed in the electrolyte containing 1.5 g/L Na3AlF6 at 125 mA/cm2 exhibited the lowest i corr of all. This suggested that at a current density of 125 mA/cm2, by adding 1.5 g/L Na3AlF6 to the electrolyte, the anticorrosion properties of the Al-Mg-Si alloy could be significantly improved with respect to the counterparts obtained in electrolytes containing either no Na3AlF6 or 3 g/L Na3AlF6. This improvement of the anticorrosion properties was attributed to the formation of a relatively defect-free structure rich in F, which prevented the diffusion of corrosive ions (Cl−) towards the substrate. In other studies, Wang et al.27 reported based on the coating layers formed on LY12 Al alloy by PEO that the i corr of the coating decreased from 1.85 × 10−7 A/cm2 to 8.8 × 10−8 A/cm2 when the electrolyte consisting of 2 g/L NaF and 8 g/L NaAlO2 was used at 80 mA/cm2 for 60 min as compared to the counterpart obtained from electrolyte without NaF. On the other hand, Arunnellaiappan et al.28 found that addition of 4 g/L CeO2 into the alkaline-silicate-based

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electrolyte would decrease the icorr from 4.2 × 10−7 A/cm2 to 1.25 × 10−10 A/ cm2 when the PEO coatings were formed on 7075 Al alloy at 150 mA/cm2 for 10 min. The present results shown in Table 3 clearly implied that the addition of 1.5 g/L Na3AlF6 into the electrolyte together with a high current density condition would beneficial for fabricating compact coatings with excellent corrosion protection properties.

Figure 5: Potentiodynamic polarization curves recorded in 3.5 wt.% NaCl solution of the coatings formed on an Al-Mg-Si alloy by PEO at (a) 50 mA/ cm2 and (b) 125 mA/cm2 using electrolytes containing different concentrations of Na3AlF6. The potentiodynamic polarization curves were recorded at a scan rate of 1 mV/s from −0.25 V to +0.4 V with respect to the open circuit potential. Table 3: Extrapolation results of the potentiodynamic polarization curves of the PEO coatings formed on an Al-Mg-Si alloy by PEO for 8 min using different concentrations of Na3AlF6 at different current densities

To investigate the effect of Na3AlF6 on the corrosion behavior of the AlMg-Si alloy coated by PEO in more detail, EIS tests in 3.5 wt.% NaCl solution were conducted, and the results are shown in Fig. 6 as the Nyquist plots. The corrosion resistance of the samples can be qualitatively compared from the EIS spectra, where larger semicircles usually indicate a higher corrosion resistance. The smallest semicircles were observed in the EIS spectrum of the coating formed with no Na3AlF6, suggesting that the corrosion resistance

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of the Al-Mg-Si alloy in 3.5 wt.% NaCl solution was improved by the coatings formed in electrolyte containing Na3AlF6. In addition, irrespective of the current density at which the coating was formed, the coating formed in the electrolyte containing 1.5 g/L Na3AlF6 exhibited the largest capacitive loop of all, indicating the excellent anticorrosion properties of this coating, which was in good agreement with the potentiodynamic polarization results shown in Fig. 5.

Figure 6: Nyquist plots recorded in 3.5 wt.% NaCl solution of the coatings formed on an Al-Mg-Si alloy by PEO at (a) 50 mA/cm2 and (b) 125 mA/cm2 using electrolytes containing different concentrations of Na3AlF6 and (c) the equivalent circuit model used for analyzing the EIS data of the coatings formed in the electrolyte containing Na3AlF6 with different concentrations at two different current densities, where R s is the resistance of the solution, R o and R i is the resistance of the outer and inner layers of the coatings, respectively, and CPE o and CPE i is the constant phase element of the outer and inner layers, respectively.

Based on the EIS results and the structure of the PEO coatings, the data was adequately fitted to the equivalent circuit model shown in Fig. 6c29. To better describe the interfacial heterogeneities of the coatings, the more general constant phase element (CPE) was used instead of a rigid capacitive element30. The CPE is defined by the following equation:

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Modern Techniques of Coating and Drying in Surface Engineering

(5) where j is the imaginary unit, ω is the angular frequency, and n and Yare the CPE parameters. The n values range from 0 to 1: for n = 0, the CPE describes an ideal resistor, and for n = 1, the CPE describes an ideal capacitor. The fitted R o , R i , n o , and n i values for the coatings obtained at different current densities using various concentrations of Na3AlF6 are presented in Fig. 7 (R is the resistance, and subscripts “o” and “i” denote the outer and inner layers of the coating, respectively).

Figure 7: Representation of the EIS-fitted R 0 , R i , n 0 , and n i values of the coatings formed at two different current densities in electrolytes containing different concentrations of Na3AlF6.

In terms of R o , which was inversely proportional to coating porosity, it was observed in Fig. 7a that the coating obtained with 1.5 g/L Na3AlF6 at 125 mA/cm2 exhibited the highest R o value, while the coating obtained with no Na3AlF6 at 50 mA/cm2 had the lowest, further confirming that in

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the formation of coatings by PEO, the addition of 1.5 g/L Na3AlF6 to the electrolyte, as well as applying a higher current density, can improve the anticorrosion properties by significantly decreasing the level of porosity. These results were consistent with the SEM results shown in Fig. 1 and Table 1. In addition, the results shown in Fig. 7b indicated that R i was generally higher than R o , indicating that the inner layer contributed the most to the overall corrosion resistance. It was also observed that the R i values for the coatings obtained in electrolytes containing Na3AlF6 were higher than for those obtained with no Na3AlF6, suggesting that Na3AlF6 also produced an increase in the compactness of the PEO coatings through the incorporation of F-containing compounds. Furthermore, CPE-n o and CPE-n i were used as an indicator to confirm microstructural results, such as porosity and micropore size, which strongly affect the anticorrosion properties of PEO coatings31, 32. As shown in Fig. 7c and d, the CPE-n o and CPE-n i values for the coatings obtained in electrolyte containing 1.5 g/L Na3AlF6 at 125 mA/cm2 were higher than for the other coatings, which indicated that the coating/electrolyte interface and coating/substrate interface became smoother with the addition of 1.5 g/L Na3AlF6 and upon increasing the current density for the PEO process.

Mechanism Underlying a Role of Na3AlF6 The effect of adding Na3AlF6 to the alkaline phosphate-based electrolyte on the coating morphology and corrosion resistance was found to be significant. During the PEO process, the AlF6 3− ions generated due to ionization of Na3AlF6 were hydrolyzed, which resulted in the formation of Al(OH)3 and F− ions (Equation 2). The high temperature under plasma conditions caused dehydration of Al(OH)3, which resulted in the formation of Al2O3 within the coating (Equations 3and 4)33. Moreover, irrespective of current density, the coatings formed in electrolyte containing 3 g/L Na3AlF6 exhibited a higher F content than their counterparts obtained with 1.5 g/L Na3AlF6, which suggested that a further hydrolysis of AlF6 3− ions occurred with 3 g/L Na3AlF6. It is worth mentioning that the presence of numerous structural defects (Figs 1a and d and 2a and d) in the coatings obtained in the electrolyte with no Na3AlF6 could provide a short path for F− ions to cross the PEO coating and distribute almost uniformly throughout it under plasma conditions (Fig. 3g). As described earlier, F− ions, being the ion with the smallest size among all negative ions, could enter the micropores, changing

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the structure and properties of the PEO coatings27. Therefore, in comparison, the coatings obtained in electrolyte containing Na3AlF6exhibited a relatively defect-free structure. However, compared with the coating obtained with 1.5 g/L Na3AlF6, the coating obtained with 3 g/L Na3AlF6, despite having the highest F content, showed larger micropores and a higher degree of porosity, which was attributed to the strong corrosive effect of the electrolyte due to formation of HF34. Indeed, a high concentration of HF could lead to the formation of numerous defects in the coating due to a decrease of the electrolyte pH, which produced etching of the substrate surface. In addition, for the coating formed in electrolyte containing 3 g/L Na3AlF6, some cracks were also observed in the higher magnification (inset of Fig. 1f), which was associated with the transformation of Al(OH)3 into Al2O3. In summary, it was concluded that Na3AlF6 could lead to an increase in the compactness of the inner layer up to a certain level, above which compactness decreased.

CONCLUSIONS A nearly defect-free coating formed by plasma electrolytic oxidation at high current density of 125 mA/cm2 was obtained in a phosphate-based electrolyte with Na3AlF6, and the role of Na3AlF6 in improving the corrosion performance of the coating was investigated. The microstructural results revealed that the coating formed in the electrolyte containing 1.5 g/L Na3AlF6 at 125 mA/cm2 exhibited the structural defects to the less extent in comparison to the other coatings since fluorine was distributed uniformly throughout the coating due to the hydrolysis of AlF6 3− triggered by intense plasma sparks. Hence, the coating formed in the electrolyte containing Na3AlF6 would improve the corrosion performance of Al-based alloys.

MATERIALS AND METHODS The substrate sample used in this study was an Al-Mg-Si alloy with a chemical composition (expressed in wt.%) of 0.99 Mg, 0.59 Si, 0.25 Cu, 0.16 Fe, 0.11 Cr, 0.13 Mn, and balance Al. Before the PEO treatment, the initial samples, with a dimension of 25 mm × 20 mm × 4 mm, were ground to 1200 grit with SiC paper and then, ultrasonically cleaned in acetone. The electrolytes were prepared by mixing 3 g/L potassium hydroxide (KOH) and 6 g/L sodium phosphate (Na3PO4) in distilled water, followed by the addition of Na3AlF6 at two different concentrations: 1.5 g/L and 3 g/L.

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The Na3AlF6 concentration was set at a maximum value of 3 g/L because higher concentrations could cause severe electrochemical etching of the substrate surface. For reference, an electrolyte containing no Na3AlF6 was also prepared. To obtain the coating, the PEO treatment was conducted at a frequency of 60 Hz during 8 min. In order to study the effect of current density on the characteristics of the coating, two different current densities, viz. 50 and 125 mA/cm2, were applied. During the PEO process, the temperature of the electrolyte was controlled to be below 25 °C by adjusting the flow rate of the cooling water. The surface morphology and elemental composition of the coating layers was studied with a scanning electron microscope (SEM, Hitachi S-4800) equipped with an energy dispersive spectrometer (EDS). The Image Analyzer 1.33 program was used to measure and calculate the micropore size and the percentage of porosity. The phase composition was examined by X-ray diffraction (XRD, Rigaku D/Max-2500) with a step size of 0.05° over a scan range from 20° to 90°. The corrosion performance of the PEO coatings was evaluated through potentiodynamic polarization and EIS (electrochemical impedance spectroscopy) tests in a 3.5 wt.% NaCl solution using a Reference 600 potentiostat from Gamry Instruments. A standard three-electrode cell with a coated alloy sample as the working electrode, Ag/AgCl (sat. KCl) as the reference electrode, and a platinum plate as the counter electrode was used in the experiments. Polarization curves were recorded from −0.25 V to +0.4 V with respect to the open circuit potential at a scan rate of 1 mV/s. Impedance measurements were performed on the PEO coatings in the frequency range from 106 Hz to 0.1 Hz with an amplitude of ±10 mV. In order to stabilize the open circuit potential, all electrochemical tests were performed after a 5-h immersion in the 3.5 wt.% NaCl solution. In addition, all tests were repeated at least three times to ensure data accuracy.

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REFERENCES 1.

Yao, Z. et al. Investigation of absorptance and emissivity of thermal control coatings on Mg–Li alloys and OES analysis during PEO process. Sci. Rep. 6, 29563, doi:10.1038/srep29563 (2016). 2. Mori, Y., Koshi, A. & Liao, J. Corrosion resistance of plasma electrolytic oxidation layer of a non-ignitable Mg-Al-Mn-Ca magnesium alloy. Corros. Sci. 104, 207–216, doi:10.1016/j.corsci.2015.12.013 (2016). 3. Kamil, M. P., Kassem, M. & Ko, Y. G. Soft plasma electrolysis with complex ions for optimizing electrochemical performance. Sci. Rep. 7, 44458, doi:10.1038/srep44458 (2017). 4. Dong, K., Song, Y., Shan, D. & Han, E. H. Corrosion behavior of a self-sealing pore micro-arc oxidation film on AM60 magnesium alloy. Corros. Sci. 100, 275–283, doi:10.1016/j.corsci.2015.08.004(2015). 5. Zhao, J., Xie, X. & Zhang, C. Effect of the graphene oxide additive on the corrosion resistance of theplasma electrolytic oxidation coating of the AZ31 magnesium alloy. Corros. Sci. 114, 146–155, doi:10.1016/j. corsci.2016.11.007 (2016). 6. Lu, X., Blawert, C., Zheludkevich, M. L. & Kainer, K. U. Insights into plasma electrolytic oxidation treatment with particle addition. Corros. Sci. 101, 201–207, doi:10.1016/j.corsci.2015.09.016 (2015). 7. Yerokhin, A. L., Nie, X., Leyland, A., Matthews, A. & Dowey, S. J. Plasma electrolysis for surface engineering: Review. Surf. Coat. Technol. 122, 73–93, doi:10.1016/S0257-8972(99)00441-7 (1998). 8. Sankara Narayanan, T. S. N., Park, S. & Lee, M. Strategies to improve the corrosion resistance of microarc oxidation (MAO) coated magnesium alloys for degradable implants: prospects and challenges. Prog. Mater. Sci. 60, 1–71, doi:10.1016/j.pmatsci.2013.08.002 (2014). 9. Malayoglu, U., Tekin, K. C. & Shrestha, S. Influence of post-treatment on the corrosion resistance of PEO coated AM50B and AM60B Mg alloys. Surf. Coat. Technol. 205, 1793–1798, doi:10.1016/j. surfcoat.2010.08.022 (2010). 10. Kassem, M., Lee, Y. H. & Ko, Y. G. Incorporation of MoO2 and ZrO2particles into the oxide film formed on 7075 Al alloy via micro-arc oxidation. Mater. Lett. 182, 260–263, doi:10.1016/j.matlet.2016.07.009 (2016). 11. Rapheal, G., Kumar, S., Scharnagl, N. & Blawert, C. Effect of current

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density on the microstructure and corrosion properties of plasma electrolytic oxidation (PEO) coatings on AM50 Mg alloy produced in an electrolyte containing clay additives. Surf. Coat.Technol. 289, 150–164, doi:10.1016/j.surfcoat.2016.01.033(2016). Lu., X., Blawert, C., Scharnagl, N. & Kainer, K. U. Influence of incorporating Si3N4 particles into the oxide layer produced by plasma electrolytic oxidation on AM50Mg alloy on coating morphology and corrosion properties. J. Magnes. Alloys. 1, 267–274, doi:10.1016/j. jma.2013.11.001 (2013). Kazanski, B., Kossenko, A., Zinigrad, M. & Lugovkoy, A. Fluoride ions as modifiers of the oxide layer produced by plasma electrolytic oxidation on AZ91D magnesium alloy. Appl. Surf. Sci.287, 461–466, doi:10.1016/j.apsusc.2013.09.180 (2013). Wang, L., Chen, L., Yan, Z., Wang, H. & Peng, J. Effect of potassium fluoride on structure and corrosion resistance of plasma electrolytic oxidation films formed on AZ31 magnesium alloy. J. Alloys Compd. 480, 469–474, doi:10.1016/j.jallcom.2009.01.102(2009). Ryu, H. S. & Hong, S. H. Effects of KF, NaOH, and KOH Electrolytes on Properties of Microarc-Oxidized Coatings on AZ91D Magnesium Alloy. J. Electrochem. Soc. 156, C298–C303, doi:10.1149/1.3158552 (2009). Liang, J. et al. Effect of potassium fluoride in electrolytic solution on the structure and properties of microarc oxidation coatings on magnesium alloy. Appl. Surf. Sci. 252, 345–351, doi:10.1016/j.apsusc.2005.01.007 (2005). Rehman, Z. R. & Koo, B. H. Combined effect of long processing time and Na2SiF6 on the properties of PEO Coatings formed on AZ91D. J. Mater. Eng. Perform. 25, 3531–3537, doi:10.1007/s11665-016-2177-2 (2016). Duan, H., Yan, C. & Wang, F. Effect of electrolyte additives on performance of plasma electrolytic oxidation films formed on magnesium alloy AZ91D. Electrochim. Acta. 52, 3785–3793, doi:10.1016/j.electacta.2006.10.066 (2007). Yerokhin, A. L., Lyubimov, V. V. & Ashitkov, R. V. Phase formation in ceramic coatings during plasma electrolytic oxidation of aluminum alloys. Ceram. Int. 24, 1–6, doi:10.1016/S0272-8842(96)00067-3 (1998).

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20. Srinivasan, P. B., Liang, J., Blawert, C., Stormer, M. & Dietzel, W. Effect of current density on the microstructure and corrosion behavior of plasma electrolytic oxidation treated AM50 magnesium alloy. Appl. Surf. Sci. 255, 4212–4218, doi:10.1016/j.apsusc.2008.11.008 (2009). 21. Yue, Y. & Hua, W. Effect of current density on corrosion resistance of micro-arc oxide coatings on magnesium alloy. Trans. Nonferrous Met. Soc. China. 20, s688–s692, doi:10.1016/S1003-6326(10)60563-8 (2010). 22. Gao, Y., Yerokhin, A. L. & Matthews, A. DC plasma electrolytic oxidation of biodegradable cp-Mg: In-vitro corrosion studies. Surf. Coat. Technol. 234, 132–142, doi:10.1016/j.surfcoat.2012.11.035(2013). 23. Zhang, X. L., Jiang, Z. H., Yao, Z. P. & Wu, Z. D. Electrochemical study of growth behavior of plasma electyrolytic oxidation coating on Ti6Al4V: Effects of the additive. Corros. Sci. 52, 3465–3473, doi:10.1016/j.corsci.2010.06.017 (2010). 24. Kaseem, M., Kamil, M. P., Kwon, J. H. & Ko, Y. G. Effect of sodium benzoate on corrosion behavior of 6061 Al alloy processed by plasma electrolytic oxidation. Surf. Coat.Technol. 283, 268–273, doi:10.1016/j. surfcoat.2015.11.006 (2015). 25. Dehnavi, V., Liu, X. Y., Luan, B. L., Shoesmith, D. W. & Rohani, S. Phase transformation in plasma electrolytic oxidation coatings on 6061 aluminum alloy. Surf. Coat. Technol. 251, 106–114, doi:10.1016/j. surfcoat.2014.04.010 (2014). 26. Malayoglu, U., Tekin, K. C., Malayoglu, U. & Shrestha, S. An investigation into the mechanical and tribological properties of plasma electrolytic oxidation and hard-anodized coatings on 6082 aluminum alloy. Mater. Sci. Eng. A. 528, 7451–7460, doi:10.1016/j. msea.2011.06.032 (2011). 27. Wang, Z., Wu, L., Cai, W., Shan, A. & Jiang, Z. Effects of fluoride on the structure and properties of microarc oxidation coating on aluminum alloy. J. Alloys. Compd. 505, 188–193, doi:10.1016/j. jallcom.2010.06.027 (2010). 28. Arunnellaiappan, T., Ashfaq, M., Krishna, L. R. & Rameshbabu, N. Fabrication of corrosion-resistant Al2O3-CeO2 composite coating on AA7075 via plasma electrolytic oxidation coupled with electrolphoretic deposition. Ceram. Int. 42, 5897–5905, doi:10.1016/j. ceramint.2015.12.136 (2016).

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29. Liu, Y., Wei, Z., Yang, F. & Zhang, Z. Environmental friendly anodizing of AZ91D magnesium alloy in alkaline borate–benzoate electrolyte. J. Alloys Compd. 509, 6440–6446, doi:10.1016/j.jallcom.2011.03.083 (2011). 30. Yao, Z., Jiang, Z., Zin, S., Sun, X. & Wu, X. Electrochemical impedance spectroscopy of ceramic coatings on Ti-6Al-4V by microplasma oxidation. Electrochim. Acta. 50, 3273–3279, doi:10.1016/j. electacta.2004.12.001 (2005). 31. Kaseem, M. & Ko, Y. G. Electrochemical response of Al2O3-MoO2-TiO2 oxide films formed on 6061 Al alloy by plasma electrolytic oxidation. J. Electrochem. Soc. 163, C587–C592, doi:10.1149/2.0931609jes (2016). 32. He, D. et al. Effect mechanism of ultrasound on growth of micro-arc oxidation coatings on A96061 aluminum alloy. Vacuum. 107, 99–102, doi:10.1016/j.vacuum.2014.04.015 (2014). 33. Al Bosta, M. M., Ma, K. & Chien, H. Effect of anodic current density on characteristics and low temperature IR emissivity of ceramic coating on aluminum 6061 alloy prepared by microarc oxidation. J. Ceram. 2013 (2013). 34. Gnedenkov, S. V. et al. Production of hard and heat-resistant coatings on aluminum using a plasma micro-discharge. Surf. Coat.Technol. 123, 24–28, doi:10.1016/S0257-8972(99)00421-1(2000).

SECTION II: COATING AND SURFACE DEFECTS

Chapter 4

Carbon based DLC films: Influence of the Processing Parameters on the Structure and Properties Francisco Andrés Delfín 1, Sonia Patricia Brühl 1, Christian Forsich 2, and Daniel Heim 2 Surface Engineering Group (GIS) – UTN FRCU – Ing. Pereyra 676, Concepción del Uruguay, Argentina. CP: 3260 1

Materials Department – FH-OOE, Wels Campus – Stelzhamerstrasse 23, Wels, Austria. CP: 4600 2

ABSTRACT Hydrogenated carbon-based films, such as DLC (“Diamond Like Carbon”), have interesting properties such as excellent tribological behavior, low friction coefficient, high superficial hardness and good wear resistance; they are chemically inert and highly corrosion resistant. They are deposited by means of PACVD (plasma-assisted chemical vapor deposition) with variable film thickness. The load carrying capacity grows with the thickness, so it

Citation: Delfín, Francisco Andrés, Brühl, Sonia Patricia, Forsich, Christian, & Heim, Daniel. (2018). “Carbon based DLC films: Influence of the processing parameters on the structure and properties”. Matéria (Rio de Janeiro), 23(2), e12059. Epub July 19, 2018. https://dx.doi.org/10.1590/s1517-707620180002.0395 Copyright © The Author(s) 2018. This is an Open Access article distributed under the terms of the Creative Commons Attribution License, which permits unrestricted use, distribution, and reproduction in any medium, provided the original work is properly cited.

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is possible to deposit thick films on “soft” steels (e.g. low alloyed steels). When increasing coating thickness, surface defects are generated during the deposition process compromising their excellent properties. In this work, different metal substrates have been used to compare adhesion and quantify superficial defects: AISI 316L, DIN 42CrMo4 (AISI 4140) and Böhler K110 (AISI D2). The films were deposited at different temperatures, changing the silicon content and the coating thickness. The samples were placed in the furnace on different positions (standing, lying or up-side down). The films were analyzed with optical and electron microscopy, 3D topography profilometer, and they were tested under sliding wear conditions. Friction coefficient and wear volume were measured, with an average friction coefficient which resulted below 0.05. A higher amount of surface defects was obtained on lying samples compared to the ones up-side down. The quantity of defects increased with the thickness of the coating and decreased with the temperature. The geometry and the growth mechanism of the defects were analyzed. Keywords Protective coatings, PACVD, DLC, surface defects

INTRODUCTION Hydrogenated amorphous carbon films (a-C:H) are a composed of carbon and hydrogen, with different C-C and C-H bondings, sp2 (graphite) and sp3 (diamond). The result is a hard coating, with extremely low friction coefficient and good wear resistance. Besides, these films are chemically inert in most aggressive environments and provide a good corrosion resistance making them attractive for many technological applications: mechanicals, electronical, biomed, and petrochemical and food industry [1-4]. As a drawback, these films have high internal stresses that result in low adhesion. The incorporation of silicon as coating dopant and an amorphous silicon interlayer helps to solve this problem, allowing to increase the load bearing capacity of the system; this coating is known as a-C:H:Si. Also, a higher amount of hydrogen and a drop of sp3 carbon bonds lead to a decrease in the film stress and helps prevent delamination [4-6]. One way of depositing these films is PACVD, using acetylene and HMDSO (hexamethyldisiloxane) as carbon and silicon precursors respectively, obtaining average deposition rates of 1 µm/h. Some authors report that they have been able to get layers up to 58 µm [6].

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The advantage of the PACVD method presented here is the use of a DC plasma nitriding equipment adapted to enable a duplex process (ion nitriding and DLC deposition at once). [6-7]. However, the DLC coatings have some imperfections in the coating that negatively influence their properties, mainly in corrosion resistance and chemical inertness. These defects enable the contact between the film surface and the substrate, allowing corrosive substances to attack the substrate and produce crevice corrosion. Some papers related to defects in DLC films have been found, but no systematic analysis which provides an insight to characterize and eventually, determine its occurrence and growing mechanism has been made up to now [8-10].

MATERIALS AND METHODS In this work, three different types of steels have been used as substrate: low alloy steel DIN 42CrMo4 (similar to AISI 4140), cold work steel Böhler K110 (AISI D2 like) and austenitic stainless steel AISI 316L. Samples in disk shape have been prepared (see Table 1), polished with Tegra System from Struers, using cloths with diamond solution of 3 µm. Table 1: Description of the samples used in this study STEEL

DENOMINATION

DIAMETER

THICKNESS

HARDNESS

DIN 42CrMo4

C

40 mm

10 mm

30 HRc

K110

A

36 mm

10 mm

60 HRc

AISI316L

SS

25 mm

5,5 mm

20 HRc

PACVD Deposition Samples were coated with a-C:H:Si using the PACVD technique (Plasma Assisted Chemical Vapor Deposition), in a commercially available ion nitriding plant, provided by the company Rübig, with a DC pulsed discharge that generates and sustains the plasma. Acetylene was used as carbon precursor and silicon was added using HMDSO. First, a silicon interlayer was deposited with a thickness of about 300 nm which is used as base layer for the deposition of the DLC film. In order to control the incidence of different parameters, several processes have been carried out by changing systematically only one parameter at a time: process temperature, silicon content and deposition time. The last one

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is correlated to the thickness of the layer, with an average deposition rate of 1 µm/h. Table 2: PACVD deposition processes PROCESS DLC-Thick DLC-Thin Middle High Temp. Low Temp. High Silicon No Silicon

TIME 35 h 3h 15 h 35 h 35 h 35 h 35 h

TEMPERATURE 450°C 450°C 450°C 550°C 400°C 450°C 450°C

HMDSO (%) 0.5% 0.5% 0.5% 0.5% 0.5% 1% 0%

For every process, four samples of each steel grade have been placed inside the deposition chamber in different positions related to the cathode plate: lying, standing and up-side down.

Coating Thickness It was measured using a Calotest [11], with a 25.4 mm steel ball.

Defects Quantification Pictures of five random zones on each sample have been taken using an optical microscope, allowing defects to be observed as black dots in the picture. Defects were counted and measured using image analysis software, by which it was possible to get a classification by size and compare the amount of defects and their size in each process condition.

Topographical Analysis A confocal microscope was used to measure different types of defects in the coatings, enabling to obtain topographical three dimensional images of the surface. To complete the analysis, a Tescan Vega SEM was used to get defects pictures. Also an EDX detector was used to analyze the chemical composition of the defects, helping to determine if the defect goes down to the substrate or not.

Growing Mechanism of Defects By means of a SEM microscope equipped with FIB (Focused Ion Beam),

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Dual Beam Helios Nano Lab, the cross section of some defects was analyzed, in order to study their sub-superficial morphology. Also in this case the EDX was used to perform the chemical analysis in the cross section.

Tribological Behavior The friction coefficient and wear rate were measured using a TRIBOtechnic Tribometer, with 12 N load, 3000 m track length and a 100Cr6 steel ball as counterpart. Once the steady-state friction coefficient was obtained, the track was measured using confocal microscopy and wear results were analyzed using specialized software.

Adhesion The adhesion strength of the deposited coatings was measured with a CSM REVETEST scratch tester. The scratch was 10 mm long, a Rockwell diamond indenter (200 µm tip radius) was used, with an increasing normal load of 100 N/min. The critical load was defined as the load at which complete delamination of the coating occurred.

RESULTS AND DISCUSSION Coating Thickness In Table Nº3 the coating thicknesses are shown, for each process group of conditions, measured by Calotest. Table 3: Coating thickness PROCESS DLC-Thick DLC-Thin Middle High Temp. Low Temp. High Silicon No Silicon

TIME 35 h 3h 15 h 35 h 35 h 35 h 35 h

TEMPERATURE 450°C 450°C 450°C 550°C 400°C 450°C 450°C

HMDSO (%) 0.5% 0.5% 0.5% 0.5% 0.5% 1% 0%

THICKNESS 40 µm 4 µm 20 µm 41 µm 38 µm 38 µm 42 µm

The calculated mean deposition rate is slightly higher than 1 µm/h, within the expected ranks. Coating thickness increases with temperature,

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whereas a rise in the silicon content in the gas mixture decreases the film thickness.

Defects Quantification The amount of defects has been counted in each sample through micrographs obtained with an optical microscope. In this section, the comparative data are displayed, describing the parameters that had more influence in the amount of defects. In Figure 1, it is possible to observe the variation in density of defects for different positions of the samples inside the deposition chamber. The result of the DIN 42CrMo4 samples with thick DLC film is shown, but the same behavior was detected in all other processes.

Figure 1: Density of defects in the samples coated with thick DLC over DIN 42CrMo4 on different positions inside the chamber.

In Figure 2 representative micrographs of the surface of each sample are shown according to their position inside the chamber. The most and bigger defects were produced on the lying sample (Figure 2a), less of them on the up-side down sample (Figure 2c).

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Figure 2: Optical images of different samples with thick coating over DIN 42CrMo4 in different positions inside the chamber. (a) Lying, (b) Vertical and (c) Up-side down.

Another parameter that influences defect formation is the process temperature, as shown in Figure 3 (a). When the process temperature increases, smaller and less defects are obtained on the surface. There is also a correlation between the amount and size of defects with the thickness of the coating, as shown in Figure 3 (b). More and bigger defects were observed when the coating was thicker. The same behavior was also found for the steel grades K110 and AISI 316L.

Figure 3: (a) Density of defects for coatings with different process temperature over DIN 42CrMo4. (b) Density of defects for different coating thickness over DIN 42CrMo4.

The reason for the defect formation seems to be dependent on the electrical conductivity of the coating. More defects are found when the conductivity is lower, caused by a thicker film or a lower process temperature. A DC unipolar pulsed electrical discharge was used, that requires a connection between the substrate (cathode) and the anode. As the coating has less conductivity than the substrate, the film starts to grow, creating some hot

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spots where the plasma concentrates and allows good conductivity, but this causes a defect in each spot.

Morphological Analysis of Defects A SEM analysis revealed the presence of different types of defects on the coating surface. The most common defect by far is the protuberance, as it can be seen in Figure 4 (a), they even got stuck together to perform much bigger defects or clusters as shown in Figure 4 (b). An EDX analysis of these defects revealed that their chemical composition is the same as the coating. Holes were also found on the surface, as it can be seen in Figure 4 (c), and it was determined that some of them could even reach the substrate. Inside these defects, some impurities have been found using EDX.

Figure 4: Different defect morphologies for thick DLC coatings on DIN 42CrMo4.

Defect Growing Mechanism By means of FIB, it was possible to analyze the cross section of the thin coatings (4 µm), where three kinds of defect growing mechanisms were found. In the first case, shown in Figure 5 (a), the coating follows the surface imperfections, such as porosity, grinding, sandblasting, etc.

Figure 5: Sub-superficial morphology of different types of defects. (a) Thin

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coating on DIN 42CrMo4. (b) Thin coating on Böhler K110 and (c) Thin coating on nitrided DIN 42CrMo4.

In Figure 5 (b) a second case is shown, where tiny dust particles are deposited on the substrate and act further as nucleation points for the film growing. Finally a third type of defect was detected, shown in Figure 5 (c), where pores have been formed near some impurities whereas the silicon interlayer seems to be unaffected. An EDX analysis was performed in the FIB cross section, revealing that the chemical composition of the coating does not seem to be affected in the first two cases (Figure 5 a and b), while in the third case (Figure 5 c) some impurities such as calcium and aluminum have been found. The subsuperficial morphology of these kinds of defects and the impurities found inside them lead to the hypothesis, that these defects become deep holes after an external force which cuts them off, looking like the holes found in the SEM analysis of the surface (Figure 4 c). It can also be considered that something similar happened to the protuberances in Figure 5 (b), due to the weakness in grip between the coating and the nucleated part that become the protuberance.

Tribology Pin-on-Disk test were carried out on all samples covered with the different coatings. No difference was observed in the tribological behavior when the substrate and the coating process parameters were changed, except for the silicon content variation that produced a change in the friction coefficient and the wear rate, as it is shown in Figure 6.

(a)

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(b) Figure 6: Tribological results for the DIN 42CrMo4 samples. (a) Friction coefficient. (b) Wear rate.

The average friction coefficient was 0.035 regardless the steel grade or the pre-treatment of the substrate. The average wear rate on the sample was 0.75x10-6 mm3/Nm.

Adhesion Scratch tests were performed on samples with different thickness. As expected, the load carrying capacity increases with the film thickness. The critical load of the coating linearly rises with increasing film thickness. In Figure 7, the experimental points together with the fitting curves are shown.

Figure 7: Critical load vs. coating thickness. Substrate: DIN 42CrMo4.

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It is possible to enhance the load carrying capacity by performing a hardening pre-treatment such as ion nitriding before the coating deposition, a process better known as duplex treatment. The nitriding pre-treatment provides a hardness gradient in the substrate that improves the interface substrate-interlayer-coating. As it is also shown in Figure 7, for all film thicknesses, the critical load was higher.

CONCLUSIONS There is a clear correlation between the position of the sample inside the process chamber and the amount of defects that appear on the sample surface, being the best option the up-side down position, which yielded the smaller density of defects. A process at the lowest temperature generated more and bigger defects, which diminish when temperature grows up. Finally, the thin coatings turned out to have less quantity of defects than the thick coatings. A hypothesis that can be proposed is about the relation between the growth of defects and the conductivity of the coating, which increases with the thickness (when the film grows) and diminishes with temperature, i.e. the higher the temperature, the higher the conductivity. The most common morphology of the defects found on the DLC surface is the protuberance, which can also form groups like protuberance islands or clusters. Some holes going down to the substrate have also been found and they can be caused by the removal of these protuberances. Deposition parameters, presence of defects and film thickness do not affect the coefficient of friction nor was the wear rate, except for the case of the silicon content, which can reduce the coefficient of friction adding 0.5%, but at the same time, wear rate slightly increased. Adhesion resulted better with the increase of thickness and the use of nitriding as pre-treatment.

ACKNOWLEDGMENTS To the Upper Austria University of Applied Sciences, Wels Campus, and the Materials Department. To the Functional Materials Department of Saarland University, Germany. To the OeAD, Austrian Agency for International Cooperation in Education and Research, for the scholarship of F. Delfin that made this research possible.

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REFERENCES 1.

DONNET, C., ERDEMIR, A., Tribology of Diamond-Like Carbon Films: Fundamentals and Applications, 1 ed., New York, Springer, 2008. 2. ROBERTSON, J., “Diamond-like amorphous carbon”, Material Science and Engineering, v. 37, n. 2-4, pp. 129-281, May 2002. 3. DALIBON, E.L., HEIM, D., et al., “Characterization of thick and soft DLC coatings deposited on plasma nitrided austenitic stainless steel”, Diamond & Related Materials, v. 59, pp. 73-79, Oct. 2015. 4. GASCO OWENS, A., BRÜHL, S.P., et al., “Comparison of Tribological Properties of Stainless Steel with Hard and Soft DLC Coatings”, Procedia Materials Science, v. 9, pp. 246-253, Apr. 2015. 5. CEMIN, F., BOEIRA, C.D., FIGUEROA, C.A., “On the understanding of the silicon-containing adhesion interlayer in DLC deposited on steel”, Tribology International, v. 94, pp. 464–469, Feb. 2016. 6. FORSICH, C., DIPOLT, C., et al., “Potential of thick a-C:H:Si films as substitute for chromium plating”, Surface & Coatings Technology, v. 241, pp. 86-9, Feb. 2014. 7. FORSICH, C., HEIM, D., MUELLER, T., “Influence of the deposition temperature on mechanical and tribological properties of a-C:H:Si coatings on nitrided and postoxidized steel deposited by DC-PACVD”, Surface & Coatings Technology, v. 203, n. 5-7, pp. 521-525, Dec. 2008. 8. BERNLAND, K., KÖHLER, B., et al., “Combined FIB technique with acoustic microscopy to detect steel–DLC interface defects”, Diamond & Related Materials, v. 15, n. 9, pp. 1405-1411, Sep. 2006. 9. HADINATA, S.S., LEE, M.T, et al., “Electrochemical performances of diamond-like carbon coatings on carbon steel, stainless steel, and brass”, Thin Solid Films, v. 529, pp. 412-416, Feb. 2013. 10. YATSUZUKA, M., TATEIWA, J., UCHIDA, H., “Evaluation of pinhole defect in DLC film prepared by hybrid process of plasmabased ion implantation and deposition”, Vacuum, v. 80, n. 11-12, pp. 1351-1355, Sep. 2006. 11. HERNANDEZ, L.C., PONCE, L., et al., “Nanohardness and Residual Stress in TiN Coatings”, Materials, v. 4, n. 12, pp. 929 – 940, May 2011.

Chapter 5

Superiority of Graphene over Polymer Coatings for Prevention of Microbially Induced Corrosion

Ajay Krishnamurthy1, Venkataramana Gadhamshetty2, Rahul Mukherjee1 , Bharath Natarajan3 , Osman Eksik1 , S. Ali Shojaee4, Don A. Lucca4, Wencai Ren5 , Hui-Ming Cheng5 & Nikhil Koratkar1,3 1 Mechanical, Aerospace and Nuclear Engineering, Rensselaer Polytechnic Institute, 110 8th Street, Troy, New York 12180, USA

Civil and Environmental Engineering, South Dakota School of Mines and Technology, Rapid City, South Dakota 57701, USA 2

3 Department of Materials Sciences and Engineering, Rensselaer Polytechnic Institute, 110 8th Street, Troy, New York 12180, USA

Mechanical and Aerospace Engineering, Oklahoma State University, 218 Engineering North, Stillwater, Oklahoma 74078, USA 4

Shenyang National Lab for Materials Science, Institute of Metal Research, Chinese Academy of Sciences, Shenyang 110016, China 5

Citation: Krishnamurthy, A. et al. “Superiority of Graphene over Polymer Coatings for Prevention of Microbially Induced Corrosion”. Sci. Rep.5, 13858 (2015). https://doi. org/10.1038/srep13858 Copyright © The Author(s) 2015. This work is licensed under a Creative Commons Attribution 4.0 International License. The images or other third party material in this article are included in the article’s Creative Commons license, unless indicated otherwise in the credit line; if the material is not included under the Creative Commons license, users will need to obtain permission from the license holder to reproduce the material. To view a copy of this license, visit http://creativecommons.org/licenses/by/4.0/

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ABSTRACT Prevention of microbially induced corrosion (MIC) is of great significance in many environmental applications. Here, we report the use of an ultra-thin, graphene skin (Gr) as a superior anti-MIC coating over two commercial polymeric coatings, Parylene-C (PA) and Polyurethane (PU). We find that Nickel (Ni) dissolution in a corrosion cell with Gr-coated Ni is an order of magnitude lower than that of PA and PU coated electrodes. Electrochemical analysis reveals that the Gr coating offers ~10 and ~100 fold improvement in MIC resistance over PU and PA coatings respectively. This finding is remarkable considering that the Gr coating (1–2 nm) is ~25 and ~4000 times thinner than the PA (40–50 nm), and PU coatings (20–80 μm), respectively. Conventional polymer coatings are either non-conformal when deposited or degrade under the action of microbial processes, while the electrochemically inert graphene coating is both resistant to microbial attack and is extremely conformal and defect-free. Finally, we provide a brief discussion regarding the effectiveness of as-grown vs. transferred graphene films for anti-MIC applications. While the as-grown graphene films are devoid of major defects, wet transfer of graphene is shown to introduce large scale defects that make it less suitable for the current application.

INTRODUCTION Literature indicates that the annual costs for corrosion1,2, including direct and indirect costs, are now approaching $1 trillion which is ~6% of the national GDP of the United States. Studies indicate that microbially induced corrosion (MIC) problems account for ~50% of the total corrosion costs3. The MIC problem spans a range of industries including aviation, oil and energy, shipping, and wastewater infrastructure1. In fact, MIC is a ubiquitous problem in the natural environment as indigenous microbes are adept at corroding metallic structures under ambient temperatures and neutral pH conditions4,5,6. MIC is caused by a genetically diverse set of microbes that exist in harmony (encapsulating themselves in a matrix of self-excreted slimy exopolymeric substance), and form a robust biological film (i.e. biofilm)3,5,7. The biofilm accelerates the corrosion process8 by modifying the chemistry of the protective metal oxide passivation layers8. Prevention of MIC is cumbersome as it requires constant detection and monitoring of microbial populations. Moreover, physical methods for eradication of biofilms (i.e. flushing) are energy-intensive and may in fact

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aggravate corrosion by dislodging oxide layers on the metal surfaces5. Metal coatings and alloys have been commercially6 used to combat corrosion in abiotic environments. However, when translated to a biotic environment their effectiveness is reduced due to aggressive microbial activity. Further, they suffer from inherent disadvantages such as environmental regulations that prohibit their use for corrosion applications (e.g. Cr)3,7,9,10. Polymer coatings (both natural and artificial) have also been used as an effective barrier for corrosion applications but can suffer from poor adhesion to the base materials and undergo rapid microbial degradation11,12,13,14,15. It has been reported that over time, pin-hole defects induced by microbial activity in polymer coatings grow in size, attract aggressive ions onto metallic surfaces, thereby further accelerating the electrochemical corrosion process16. Moreover, the typical thickness of commercial polymeric coatings17 disrupts the functionality (e.g. electrical and thermal conductivity) and dimensional tolerances of target metals. Graphene (Gr), a two-dimensional sheet of sp2 bonded carbon atoms, can be employed as an ultra-thin corrosion-resistant coating, as it is mechanically robust, flexible, chemically inert, thermally and electrically conductive, and can form an impermeable barrier18,19,20,21,22,23. Further, ultra-thin graphene coatings can be applied without negatively impacting the functionality (e.g. electrical, thermal conductivity etc.) and dimensions of the underlying metal. Such graphene coatings have been recently demonstrated as corrosion-resistant coatings for metals (e.g. Ni, Cu, Fe, and steel alloys) under abiotic environments24,25,26. However, these studies were based on relatively short time scales (minutes to hours). Recently, two studies have provided some very interesting observations on the failure of graphene coatings on copper substrates under abiotic conditions27,28. The reason for coating failure was attributed to mass transport through the nanoscale defects present on the graphene sheet, which can be reduced significantly by the use of few-layer graphene29. Further, it has been shown that defect plugging (using passive Al2O3 nanoparticles) caused a significant improvement in the corrosion resistance of monolayer graphene29. In our recent study, we found that 3–4 layer graphene films deposited by chemical vapor deposition (CVD) offer long-term resistance (~2400 h) to bimetallic corrosion of Ni, especially under microbial conditions30. In this work, we compare the MIC resistance of graphene to two widely used polymer coatings. In particular parylene (PA) is one of the most popular barrier coatings used by industry as it has excellent mechanical properties and provides pin-hole free coatings. Polyurethane (PU) is also widely used to protect surfaces. A detailed

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electrochemical analysis reveals that the graphene coating offers ~10-fold improvement in MIC resistance compared to PU and ~100-fold compared to PA. This finding is remarkable considering that the average thickness of the graphene coating (1–2 nm) is ~25-fold smaller than PA (40–50 nm), and ~4000-fold lower than the PU coating (20–80 μm). Post-mortem analysis reveals that graphene is highly resistant to microbial attack as compared to the polymer coatings. We perform detailed microbial analysis to comprehend the success of graphene coatings and the failure of polymer coatings. We also compare as-grown vs. transferred graphene films and show that transferred films are far more defective than as-grown ones. In addition, we explore the effect of number of graphene layers and show that few-layered (as-grown) graphene films offer by far the most defect-free surfaces and are therefore the best suited for microbial corrosion resistance applications. Since many of graphene’s fascinating properties strongly depend on the concentration and nature of defects, the defect analysis presented in this study may find broad use in several graphene based applications.

RESULTS AND DISCUSSIONS The objective of this study was to investigate the effectiveness of nanoscale graphene coatings to prevent MIC of the immersed, metallic structure under galvanic conditions. More specifically, the MIC resistance of Ni was compared in presence of three different coatings: i) parylene-C coating on Ni (PA/Ni), ii) polyurethane coating on Ni (PU/Ni) and iii) graphene coating on Ni (Gr/Ni). A Ni foam30 was used as the substrate for all three coatings. Chemical vapor deposition (CVD- see methods section) was used to obtain a uniform coating of PA–C on the Ni surface (Fig. 1a). The thickness of the coating was determined using ellipsometry. The ratios of the amplitude (ψ), and phase changes of the p and s components (Δ) of polarized light, are shown in (Fig. 1b). Using the Cauchy coefficients (A = 1.58 and B = 0.012)31, the average thickness of the PA film was estimated to be ~46.1 nm. The PU films were deposited using standard spray coating (see methods section). As shown in Fig. 1c,d, scanning electron microscopy (SEM) imaging was used to determine the thickness of PU/Ni in the range of 20–80 μm. The graphene films (Fig. 1e) were deposited on the Ni foam using CVD (see Methods section) as shown in our previous study30. From the Raman analysis of the Gr-Ni sample (Fig. 1f), the I(2D)/I(G) ratio30,31,32 indicates a fingerprint of few-layered (3–4 layers) graphene coatings. This was also confirmed by high-resolution scanning electron microscopy (Fig. 2a,b) and transmission

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electron microscopy (Fig. 2c,d) of the graphene coating. Further, the absence of the Raman D band (~1350 cm−1) in Fig. 1fconfirms the low density of defects in the Gr film32.

Figure 1: Dimensional characteristics of three coatings on Nickel foam surfaces: (a) Parylene (PA) coated Nickel, (b) Ellipsometry shows ~46 nm PA coating, (c) Polyurethane (PU) coated Nickel, (d) SEM image showing thickness of PU coating of 20–80 microns, (e) Conformal coating of graphene film on a Ni foam and (f) Raman spectra of Gr/Ni foam at three different locations indicating that the graphene film is on average comprised of few-layer (~3–4 layer) graphene.

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Figure 2: (a) Low magnification scanning electron microscopy (SEM) image of Ni foam, (b) high magnification SEM image, showing the wrinkled graphene topography on the Ni foam surface, (c) low magnification transmission electron microscopy (TEM) image of few-layer graphene on Ni, (d) high magnification TEM image showing 4 layers of graphene on the Ni surface.

The corrosion cells were operated in a fed-batch mode for 30 days to promote the MIC conditions that induce bimetallic corrosion of the Ni anode.

Figure 3: (a) SEM image of biofilm on Gr/Ni, (b) MIC-resistant Gr/Ni anode after 30 days of MIC testing, (c) SEM image of biofilm on PA/Ni, (d) Corroded Ni/PA anode after 30 days of MIC experiment, (e) SEM image of biofilm on PU/Ni and (f) Corroded Ni/PU anode after 30 days of MIC experiment.

The SEM micrographs (Fig. 3a,c,e) provide a clear evidence for the presence of biofilm on the surfaces of Gr/Ni, PA/Ni, and PU/Ni. Further, visual

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examination of PA/Ni (Fig. 3d) and PU/Ni (Fig. 3f) indicates the presence of green corrosion by-products (e.g. Ni (II) compounds). The edges of the PA/ Ni were found to be corroded possibly due to direct (metal-microbe interaction) and indirect action of microbes (intermediate metabolites such as volatile fatty acids)10. Figure 3suggests that the anti-MIC behavior of PU/Ni is slightly better than that of PA/Ni. The most important observation is that the graphene coating was able to preserve the underlying Ni material (Fig. 3b) even after continuous exposure to the MIC environment for 30 days.

Electrochemical Studies The Nyquist plots (Fig. 4a) shows the anodic impedance (sum of polarization resistance and electrolyte resistance) of the three corrosion cells in a complex-impedance-plane. The asymptote in the low-frequency region of the Bode plots (Fig. 4b) represents the total impedance to MIC corrosion of Nickel. Among the three coatings, graphene coating (Gr/Ni) offers the highest anodic impedance followed by PU/Ni and PA/Ni (Fig. 4a,b). Another important finding can be inferred from the locus of the points for PU/Ni and PA/Ni in Fig. 4a; both the curves do not trace a true semicircle implying that the anodic impedance response does not correspond to a single activationenergy controlled process.

Figure 4: (a) Nyquist Plots for the three coatings, (b) Bode Modulus plots for the three coatings, (c) Equivalent circuit for PU/Ni, (d) Equivalent circuit for PA/Ni and (e) Equivalent circuit for Gr/Ni.

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Figure 4a,b indicates that dissimilar anodic processes occur in Gr/Ni, PU/ Ni, and PA/Ni. Hence, we used three different equivalent circuits (Fig. 4c–e) that are descriptive of the three coatings varying in thickness and anti-MIC properties. In the PU/Ni, the thickness of PU coating ranged from 20 to 80 μm, thereby the Ni foam can be considered to be embedded in the aggregate mixture of polymer matrix and biofilm33. The aggregate film on the Ni surface is porous and the electrochemical reactions are likely to occur only on the exposed surfaces at the end of a pore (Fig. 4c). At the interface located at the end of the pore, the corresponding corrosion impedance is the parallel combination of charge transfer resistance (Rct) and interfacial capacitance (Cm). The seepage of anolyte through the PU coating is responsible for the charge transfer reactions associated with the corrosion process. In order to model diffusion limitations of corrosion by-products through the porous electrode, a Warburg element (WO) is added in series to the corrosion impedance. Within the pore length, the electrolyte resistance is Rint, and the insulating part of the coating can be considered to be a capacitor Cdl which is in parallel with the impedance in the pore (Fig. 4c). The electrolyte resistance (Rel) is added in series with the previous impedance. The thickness of the parylene coating in the PA/Ni system is ~3 orders of magnitude smaller than that of PU in PU/Ni. We therefore used an alternate equivalent circuit model based of the anolyte resistance (Rel), the anolyte/ PA interfacial impedance and PA/Ni interfacial impedance (Fig. 4d). At the anolyte/PA interface, the defects and the pores in the PA coating enable the anolyte to access the underlying nickel surface, and such impedance is modeled as a pore resistance (Rint). Similarly, the interfacial capacitance at the anolyte/PU interface is modeled as Cdl. The charge transfer reactions associated with Ni corrosion at the anolyte/PA interface were incorporated as a parallel combination of charge transfer resistance (Rct) and interfacial capacitance (Cint)34,35. We modeled MIC of Gr/Ni as a modified Randles equivalent circuit (Fig. 4e)30,36. The circuit fitting analysis revealed that the Rct of Gr/Ni (~35.8 kΩ.cm2) was ~10-fold higher than that of PU/Ni (~1.57 kΩ.cm2), and ~100-fold higher than that of PA/Ni (~0.38 kΩ.cm2). We conclude that the graphene coating offers superior MIC-corrosion resistance as compared to PA and PU coatings. The cyclic voltammetry (CV) results also corroborate that graphene provides an inert environment that suppresses Ni corrosion (Fig. 5a). The PU/Ni and PA/Ni registered a higher range of electrochemical current compared to that of Gr/Ni. The maximum current in Gr/Ni was at

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least 10,000 fold lower than that of PU/Ni. A broad shoulder was observed in the PU/Ni and the PA/Ni responses between 0 and 0.8 V (vs. Ag/AgCl). This denotes a redox reaction occurring on the Ni surface. In comparison, the Gr anode response was devoid of significant oxidation or reduction peaks.

Figure 5: (a) Cyclic Voltammograms for the three coatings in corrosion cells, (b) Soluble nickel concentration in the anolyte for the three corrosion cells. Note: Corrosion cells were operated in a fed-batch mode; temporal graph for soluble Ni will not follow a linear pattern.

Nickel Dissolution Figure 5b shows the Ni concentrations in the spent-anolyte at end of three different fed-batch cycles in the three corrosion cells. The concentration of Ni2+ in both the PA/Ni and PU/Ni cells rises significantly during the course of three weeks. The corrosion rates during Week 1 were ~2.22 mg/L/day and ~0.915 mg/L/day in PA/Ni and PU/Ni, respectively; in week 3, the rates reached as high as ~6.708 mg/L/day and ~5.229 mg/L/day, respectively. Further, the Ni concentration in the cells with polymer coatings (week 2: PU/Ni ~22.89 mg/l, PA/Ni ~36.08 mg/l, week 3: PU/Ni ~36.604 mg/l, PA/ Ni ~46.962 mg/l) increased linearly. In Gr/Ni, the corrosion rates remained

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in the 1.5–3 mg/L/day range in all the three cycles. The cell with Gr/ Ni demonstrated an improvement in the corrosion resistance (week 2: ~14.663 mg/l, week 3: ~11.938 mg/l). The higher values of Ni concentrations in PU/Ni and PA/Ni cells during subsequent weeks denote an eventual failure of the polymer coatings. The lower corrosion rates in the Gr/Ni are likely due to a combination of the following: i) minimal defects in the Gr coating compared to PA and PU coatings, ii) limited diffusion of corrosion byproducts from the Ni to the bulk liquid due to restricted access through the few-layered Gr coating, and iii) adhesion of polysaccharides and insoluble biomass to the hydrophobic graphene surface which plugs the few defect sites that may exist on the Gr film surface.

Figure 6: Coating failure by (a) Localized tears in PA/Ni electrode, (b) Nonconformity and poor adhesion in PU/Ni electrode, Raman mapping of (c)

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monolayer graphene grown on copper (bright spots indicate point defects), (d) monolayer graphene transferred onto SiO2, (e) bilayer graphene transferred onto SiO2, (f) trilayer graphene transferred onto SiO2, (g) Few-layer graphene grown in-situ on the Nickel Foam by chemical vapor deposition. This as-grown few-layered graphene film is relatively free of defects and provides excellent barrier protection. The scale bars in (c) and (g) are 6 μm and in (d–f) are 10 μm.

Scanning electron microscopy (SEM) was used to observe the extent of corrosion-induced debilitation of the PA/Ni and PU/Ni surfaces. We found micron-length tears on the PU/Ni surfaces (Fig. 6a) while the PU/ Ni surfaces exhibited poor conformity and poor adhesion to the Ni surface (Fig. 6b). The microbially induced degradation and tearing of the PU coating that we report here is consistent with what has been reported in the literature37,38,39,40. The micron-scale tears on the PU/Ni surface lead to i) enhanced charge transfer associated with Ni corrosion, and ii) increased diffusion of intermediate metabolic byproducts (e.g. acetic acid resulting from glucose fermentation) on the Ni surface. Figure 6b indicates the nonconformal nature of the PU coating, and the vacancies at the PU/Ni interface which expose the Ni surface to microbial corrosion. Figure 6a,b justify our observations on the gradual increase in Ni corrosion rates with time (high Ni corrosion rates for PA/Ni and PU/Ni appeared during the later stages of corrosion-cell operation). In comparison, the Gr coating, which is extremely conformal, provides effective barrier protection, and restricts the microbes from accessing the Ni surface, thereby supporting the observed low corrosion rates of the Gr/Ni system shown in Fig. 5b.

Defect Analysis of Graphene Coatings On having established that CVD-grown (few-layer) graphene offers superior resistance to MIC, we sought to characterize the defectiveness of the as-grown graphene coating when compared to graphene films that are transferred (wet-transfer19) onto surfaces that are incompatible with CVD growth. For this purpose, monolayer graphene, grown in-situon a copper foil by CVD, was transferred layer-by-layer onto SiO2 using wet chemistry based transfer methods19. Confocal Raman spectroscopy maps of the defect intensity ratio (I(D)/I(G)) were generated for monolayer graphene grown on Cu (~30 μm × 30 μm) (Fig. 6c) and monolayer (Fig. 6d), bilayer (Fig. 6e) and trilayer (Fig. 6f) transferred graphene sheets on SiO2 (~80 μm × 80 μm).

The mapping in Fig. 6c shows point defects (bright spots) on the Gr/ Cu surface, which can act as the initiation points for corrosion as observed

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by Schriver et al. and Zhou et al.27,28. The Raman maps for the transferred samples (Fig. 6d–f) show an increased presence of defective regions in comparison to the as-grown monolayer graphene (Fig. 6c). These results suggest that the transfer process induces a large number of defects in the graphene films. The overall I(D)/I(G) defect intensity ratio was calculated for samples and was found to be about an order of magnitude lower for the few-layer graphene grown in-situ on Ni (Gr/Ni), as compared to the Gr films transferred onto SiO2 and half the value for monolayer Gr as-grown on Cu foil. This is visually represented in the Raman map for the few-layer graphene grown on Ni foam (Gr/Ni- Fig. 6g) used in our testing (Figs 3, 4, 5). Note that the Raman map in Fig. 6g is completely devoid of any bright spots (i.e. defects). Scanning electron microscopy (SEM) images for these samples indicate large scale wrinkles in the monolayer graphene transferred onto SiO2 and a notable reduction in the wrinkle density observed for two and three layer transferred graphene samples. Note that the highest defect intensity ratio recorded for the monolayer graphene (post-transfer) can be attributed to strong adhesion between SiO2 and the graphene layer41. On adding further graphene layers, the weak interactions between the similar graphene layers allows the freshly transferred graphene sheet to slide more easily on top of the previous sheet, thereby reducing stress build-up and defect generation during the transfer process. However the few-layer graphene films that are as-grown on the metal surface are by far the least defective (see Fig. 6g) indicating that it is essential to deposit the graphene ‘in-situ’ using bottom-up growth techniques such as chemical vapor deposition. Wet-chemistry based transfer methods19 to deposit graphene coatings are unlikely to yield pin-hole free coatings. These results indicate that it is crucial to directly grow few-layered graphene films on the surfaces to be passivated rather than post-synthesis transfer onto the surface.

Microbial Phylogenetic Analysis and Mechanism for Superior Corrosion Resistance of Graphene Corrosion characteristics are determined by the corroding and the corroded species. In order to better understand the interactions between the microbial community and the coated nickel surfaces, a phylogenetic microbial community analysis was carried out and its results are presented in this section (Fig. 7a). The anaerobic conditions prevalent in the corrosion cell discourage the growth of aerobic microbial communities, thereby promoting

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anaerobic fermentation and acid production which accelerate the MIC process. Literature42,43,44indicates that Stenotrophomonas spp., previously isolated from various environments including corroded metal surfaces are capable of degrading xylene-based compounds. These studies suggest that the Stenotrophomonas spp. may be responsible for the tears observed on parylene-coated nickel surface. Coupled with the acidic environment caused by the presence of Dysognomonas spp.45, Desulfovibrio spp.46, Clostridium spp.47,48, there is rapid degradation of polymer coatings and the nickel foam surface after biofilm formation. Figure 7b depicts the participating microbial and electrochemical reactions that influence the MIC of the Ni surface. The Ni surface is hypothesized to interact with the native Ni oxide, biofilm, the bulk anolyte, and the protective coating. In the absence of coatings, the bare Ni anode undergoes MIC due to the direct and indirect influence of anaerobic bacteria such as acid producing (APB) and sulfur reducing bacteria (SRB) 46. It should be noted that the nutrients (e.g. glucose) in the anolyte support the growth of APB that secrete organic acids (e.g. acetate)49,50 (Fig. 7a), and accelerate MIC of Ni surface. Finally, the SRB accelerate MIC of Ni surface5,46,51 in the following sequence: i) SRB reduce sulfur compounds to HS− ions ii) sulfide-oxidizing bacteria convert HS− to sulfate that is quickly oxidized to sulfuric acid and abruptly decrease the local pH (100 Hz) into the which spinel framework [22]. For PMF PMF carbon coated EIS spectra reflects the contact resistances between the LiNi0.5 Mn1.5 O4 electrode, the peaks around 4 V should discharge and charge performance of lithium nickel manganese oxide 3+ active materials and electrolyte or current collector [24]. weaken obviously. This indicates that the Mn ions dissolution amount of the PMF carbon coated LiNi0.5 Mn1.5 O4 is smaller than no carbon coating [23]. The presence of Mn3+ in the LMNO material plays an important role in the capacity retention [9]. And the two redox peaks of LiNi0.5 Mn1.5 O4 with 10 wt.% PMF carbon coating are overlapped to a broad peak. The cyclic voltammetry curves of the PMF carbon coated LiNi0.5 Mn1.5 O4 material exhibit smaller potential intervals that indicated higher electrode reaction reversibility

The significant improvements in electric conductivity for LiNi0.5 Mn1.5 O4 electrode could therefore be attributed to the PMF carbon coating [25].

4. Conclusions PMF carbon coated LiNi0.5 Mn1.5 O4 was prepared as the 5 V cathode materials of lithium ion batteries for the first

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materials. Another reason might be that the PMF carbon coating prevents the Mn3+ dissolving in the process of charging and discharging, thereby reducing the fading rate of discharge capacity [20, 21].

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Figure 4 shows the 2nd, 50th, and 100th charge and discharge curves of carbon coated LiNi0.5Mn1.5O4 at the rate of 10 C. The specific discharge capacity of LiNi0.5Mn1.5O4 declines sharply after PMF carbon coating at high rate discharge. The 2nd cycle specific discharge capacity decreased from 117.9 mAh·g−1 (before coating) to 110.6 mAh·g−1 (after coating). The PMF carbon coating hinders the Li+ ionic conduction at high rate discharge and charge. But the corresponding capacity retention was greatly improved after PMF carbon coating. The 100th cycle specific discharge capacity of carbon coated LiNi0.5Mn1.5O4 was 105.4 mAh·g−1, and even the corresponding capacity retention was 95.2%. And yet the specific discharge capacity of no coated LiNi0.5Mn1.5O4 was 85.4 mAh·g−1, the corresponding capacity 4 International Journal of Polymer Science retention was 69.1%.

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Figure 4: The charge and discharge curves of prepared LiNi0.5Mn1.5O4 at 10 C.

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500 0.9 Figure 5 shows the CV curves of the PMF carbon coated LiNi0.5Mn1.5O4 Q Q electrode. In the full range view of CV400curves, one pair of redox peaks 0.6 around 4 V (Mn3+/Mn4+) and two pairs of well separated strong redox peaks 0.3 300 at 4.6–4.8 V can be observed in CV curve of no carbon coated LiNi0.5Mn1.5O4 0.0 electrode. The two strong redox peaks indicate a two-stage (Ni2+/Ni3+ and 200 3+ 4+ + Ni /Ni ) Li extraction from or insertion into the spinel framework [22]. For −0.3 100 PMF carbon coated LiNi0.5Mn1.5O4 electrode, the peaks around 4 V should −0.6 3+ weaken obviously. This indicates that the 0Mn ions dissolution amount of 0 100 200 300 400 500 3.0 3.5 4.0 4.5 5.0 Z (Ω) E (V versus Li /Li) the PMF carbon coated LiNi0.5Mn1.5O4 is smaller than no carbon coating 󳰀󳰀

+

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Figure 5: Cyclic voltammogram curves of prepared LiNi0.5 Mn1.5 O4 .

LiNi0.5 Mn1.5 O4 . The as-prepared LiNi0.5 Mn1.5 O4 has the cubic face-centered spinel structure with a space group of Fd3m. The PMF carbon coating greatly promotes electrochemical performance and reduces the LiNi0.5 Mn1.5 O4 elec-

No-PMF PMF-10%

Figure 6: EIS spectra of prepared LiNi0.5 Mn1.5 O4 .

Acknowledgments This work was supported by the National Natural Science Foundation of China (no. 21376056 and no. 21463030), China Postdoctoral Science Foundation (2016M590781),

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[23]. The presence of Mn3+ in the LMNO Capacitymaterial (mAh·g−1 ) plays an important role in Capacity the capacity retention [9]. 2nd And the two redox peaks of LiNi0.5Mn1.5O4 with2nd 10 wt.% PMF carbon coating 50th 50th are overlapped to a broad peak. The cyclic 100th 100thPMF carbon coated LiNi Mn O material voltammetry curves of the 0.5 1.5 4 (a) No-PMF (b) PMF-10 exhibit smaller potential intervals that indicated higher electrode reaction reversibility and lower polarization. carbon coating FigureThe 4: ThePMF charge and discharge curves of would prepared LiNi0.5 Mn1.5 O4 at 10 effectively restrain the Mn3+ stripping and the side reaction between positive electrode active material and electrolyte at high voltage. 500

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Figure 6: EIS spectra of p

Figure 5: Cyclic voltammogram curves of prepared LiNi0.5Mn1.5O4.

Acknowledgments Figure 6 LiNi depicts the electrochemical impedance spectroscopy of the 0.5 Mn1.5 O4 . The as-prepared LiNi0.5 Mn1.5 O4 has the electrode obtained after the first of charge. Thegroup semicircle at work high was supported by cubic face-centered spinelcycle structure with a space of This Fd3m. TheinPMF greatly electrofrequency (>100 Hz) EIS carbon spectracoating reflects the promotes contact resistances between Foundation of China (no. 2 chemical performance and reduces the LiNi0.5 Mn1.5 O4 elecChina Postdoctoral Science the active materials and electrolyte or current collector [24]. The significant trode polarization in the process of charging and discharging. and could Science and Technology improvements electric conductivity for LiNi0.5Mn O electrode improved Theincycle life of LiNi 1.5 4 by 0.5 Mn1.5 O4 was significantly 201509030005 and no. 2015100 therefore be attributed the PMF carbon coating [25]. PMF carbontocoating, especially at high rate discharging and charging.

Competing Interests The authors declare that they have no competing interests.

References

[1] R. Santhanam and B. Rambabu age spinel LiNi0.5 Mn1.5 O4 ma vol. 195, no. 17, pp. 5442–5451

Capacity (mAh·g )

Capacity (mAh·g ) 2nd 50th 100th

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(b) PMF-10%

Figure 4: The charge and discharge curves of prepared LiNi0.5 Mn1.5 O4 at 10 C.

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500 Q

Q

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400 300 200 100

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5.0

0

0

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300

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ogram curves of prepared LiNi0.5 Mn1.5 O4 .

Figure 6: EIS spectra of prepared LiNi0.5 Mn1.5 O4 .

Figure 6: EIS spectra of prepared LiNi0.5Mn1.5O4.

Acknowledgments s-prepared LiNi0.5 Mn1.5 O4 has the nel structure with a CONCLUSIONS space group of This work was supported by the National Natural Science n coating greatly promotes electroFoundation of China (no. 21376056 and no. 21463030), and reduces the LiNi0.5 Mn1.5 O4 elecChina LiNi Postdoctoral PMF carbon coated Mn1.5Science O4 wasFoundation prepared(2016M590781), as the 5 V cathode materials e process of charging and discharging. and Science 0.5 and Technology Project of Guangzhou (no. of improved lithiumbyion 201509030005 batteries for the first time. PMF carbon coating increases Mn1.5 O4 was significantly and no. 201510010131). specially at high rate discharging and the crystallinity of LiNi Mn O . The as-prepared LiNi Mn O has the 0.5

1.5

4

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1.5

4

cubic face-centered spinel structure with a space group of Fd3m. The PMF References carbon coating [1]greatly promotes electrochemical performance and reduces R. Santhanam and B. Rambabu, “Research progress in high voltsts the LiNi0.5Mn1.5Oage electrode polarization in the process spinel LiNi0.5 Mn of Power Sources,of charging and 1.5 O4 material,” Journal 4 vol. 195, no. 17, pp. 5442–5451, 2010. t they have no competing interests. discharging. The cycle life of LiNi0.5Mn1.5O4 was significantly improved by PMF carbon coating, especially at high rate discharging and charging.

ACKNOWLEDGMENTS This work was supported by the National Natural Science Foundation of China (no. 21376056 and no. 21463030), China Postdoctoral Science Foundation (2016M590781), and Science and Technology Project of Guangzhou (no. 201509030005 and no. 201510010131).

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REFERENCES 1.

R. Santhanam and B. Rambabu, “Research progress in high voltage spinel LiNi0.5Mn1.5O4 material,” Journal of Power Sources, vol. 195, no. 17, pp. 5442–5451, 2010. 2. H.-S. Fang, Z.-X. Wang, X.-H. Li, H.-J. Guo, and W.-J. Peng, “Exploration of high capacity LiNi0.5Mn1.5O4 synthesized by solidstate reaction,” Journal of Power Sources, vol. 153, no. 1, pp. 174–176, 2006. 3. Y. Qian, Y. Deng, Z. Shi, Y. Zhou, Q. Zhuang, and G. Chen, “Submicrometer-sized LiMn1.5Ni0.5O4spheres as high rate cathode materials for long-life lithium ion batteries,” Electrochemistry Communications, vol. 27, pp. 92–95, 2013. 4. J.-H. Kim, N. P. W. Pieczonka, Z. Li, Y. Wu, S. Harris, and B. R. Powell, “Understanding the capacity fading mechanism in LiNi0.5Mn1.5O4/ graphite Li-ion batteries,” Electrochimica Acta, vol. 90, pp. 556–562, 2013. 5. D. Liu, J. Han, and J. B. Goodenough, “Structure, morphology, and cathode performance of Li1-x[Ni0.5Mn1.5]O4 prepared by coprecipitation with oxalic acid,” Journal of Power Sources, vol. 195, no. 9, pp. 2918– 2923, 2010. 6. Y.-C. Jin and J.-G. Duh, “Nanostructured LiNi0.5Mn1.5O4 cathode material synthesized by polymer-assisted co-precipitation method with improved rate capability,” Materials Letters, vol. 93, no. 2, pp. 77–80, 2013. 7. L. Wang, D. Chen, J. Wang, G. Liu, W. Wu, and G. Liang, “Synthesis of LiNi0.5Mn1.5O4 cathode material with improved electrochemical performances through a modified solid-state method,” Powder Technology, vol. 292, pp. 203–209, 2016. 8. J. Hassoun, K.-S. Lee, Y.-K. Sun, and B. Scrosati, “An advanced lithium ion battery based on high performance electrode materials,” Journal of the American Chemical Society, vol. 133, no. 9, pp. 3139–3143, 2011. 9. J. Xiao, X. Chen, P. V. Sushko et al., “High-performance LiNi0.5Mn1.5O4 Spinel controlled by Mn3+concentration and site disorder,” Advanced Materials, vol. 24, no. 16, pp. 2109–2116, 2012. 10. J. Deng, Y. Xu, L. Xiong, L. Li, X. Sun, and Y. Zhang, “Improving the fast discharge performance of high-voltage LiNi0.5Mn1.5O4 spinel by Cu2+, Al3+, Ti4+ tri-doping,” Journal of Alloys and Compounds, vol.

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677, no. 8, pp. 18–26, 2016. G. Liu, Y. Du, W. Liu, and L. Wen, “Study on the action mechanism of doping transitional elements in spinel LiNi0.5Mn1.5O4,” Electrochimica Acta, vol. 209, pp. 308–314, 2016. S. Wang, P. Li, L. Shao et al., “Preparation of spinel LiNi0.5Mn1.5O4 and Cr-doped LiNi0.5Mn1.5O4cathode materials by tartaric acid assisted solgel method,” Ceramics International, vol. 41, no. 1, pp. 1347–1353, 2015. Y. S. Lee, Y. K. Sun, S. Ota, T. Miyashita, and M. Yoshio, “Preparation and characterization of nano-crystalline LiNi0.5Mn1.5O4 for 5 V cathode material by composite carbonate process,” Electrochemistry Communications, vol. 4, no. 12, pp. 989–994, 2002. H. B. Yao, Y. Xie, G. J. Han, and D. M. Jia, “Enhanced electrochemical performance of the cathode material LiNi0.5Mn1.5O4 embedded by CNTs,” Journal of Materials Science: Materials in Electronics, vol. 26, no. 3, pp. 1780–1783, 2015. X. Li, W. Guo, Y. Liu, W. He, and Z. Xiao, “Spinel LiNi0.5Mn1.5O4 as superior electrode materials for lithium-ion batteries: ionic liquid assisted synthesis and the effect of CuO coating,” Electrochimica Acta, vol. 116, no. 1, pp. 278–283, 2014. T. Hwang, J. K. Lee, J. Mun, and W. Choi, “Surface-modified carbon nanotube coating on high-voltage LiNi0.5Mn1.5O4 cathodes for lithium ion batteries,” Journal of Power Sources, vol. 322, no. 8, pp. 40–48, 2016. A. Vijn, F. Hoffmann, and M. Fröba, “Thermal conversion to form LiNi0.5Mn1.5O4−δ: Influence of precursors and supporting carbon template materials,” Thermochimica Acta, vol. 638, pp. 138–150, 2016. Y. Wei, C. Shengzhou, and L. Weiming, “Oxygen reduction on nonnoble metal electrocatalysts supported on N-doped carbon aerogel composites,” International Journal of Hydrogen Energy, vol. 37, no. 1, pp. 942–945, 2012. Y.-J. Gu, Y. Li, Y.-B. Chen, and H.-Q. Liu, “Comparison of Li/Ni antisite defects in Fd-3 m and P4332 nanostructured LiNi0.5Mn1.5O4 electrode for Li-ion batteries,” Electrochimica Acta, vol. 213, no. 9, pp. 368–374, 2016. H. Wang, Z. Shi, J. Li et al., “Direct carbon coating at high temperature

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Chapter 9

Washcoat Deposition of Ni- and Co-ZrO2 Low Surface Area Powders onto Ceramic Open-Cell Foams: Influence of Slurry Formulation and Rheology Riccardo Balzarotti, Mirko Ciurlia, Cinzia Cristiani , and Fabio Paparella Politecnico di Milano, Dipartimento di Chimica, Materiali e Ingegneria Chimica “G. Natta”, Piazza Leonardo da Vinci 32, 20133 Milano, Italy

ABSTRACT The effect of formulations and procedures to deposit thin active layers based on low surface area powders on complex geometry substrates (open-cell foams) was experimentally assessed. An acid-free liquid medium based on water, glycerol, and polyvinyl alcohol was used for powder dispersion,

Citation: Balzarotti, Riccardo; Ciurlia, Mirko; Cristiani, Cinzia; Paparella, Fabio. 2015. “Washcoat Deposition of Ni- and Co-ZrO2 Low Surface Area Powders onto Ceramic OpenCell Foams: Influence of Slurry Formulation and Rheology.” Catalysts 5, no. 4: 2271-2286. https://doi.org/10.3390/catal5042271 Copyright © 2015 by the authors; licensee MDPI, Basel, Switzerland. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution license (http://creativecommons.org/licenses/by/4.0/).

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while a dip-coating technique was chosen for washcoat deposition on 30 PPI ceramic open-cell foams. The rheological behavior was explained on the bases of both porosity and actual powder density. It was proved that the use of multiple dippings fulfills flexibility requirements for washcoat load management. Multiple depositions with intermediate flash drying steps at 350 °C were carried out. Washcoat loads in the 2.5 to 22 wt. % range were obtained. Pore clogging was seldom observed in a limited extent in samples with high loading (>20 wt. %). Adhesion, evaluated by means of accelerated stress test in ultrasound bath, pointed out good results of all the deposited layers. Keywords: ceramic open-cell foams, washcoat, catalyst deposition, rheology, structured support

INTRODUCTION Structured catalysts and reactors for process intensification are receiving a large interest from the modern chemical engineering community [1]. A structured catalyst is intended as a continuum metallic or ceramic geometrical matrix (support) where voids (i.e., channels) are present. The catalytically active phases are properly dispersed onto the support surface by direct incorporation or coating deposition. The latter is the most used technique, due to its simplicity and versatility [2]. Commonly, the deposited washcoat consists of a high surface area carrier, generally ceramic, where the metal active phase is properly dispersed. A variety of materials have been investigated as high surface area ceramic supports, such as alumina, silica, titania, and ceria [3,4,5,6]. In some process applications, low surface area catalysts have been investigated instead of high surface area ones. As an example, cerium oxide has been proposed as an active phase carrier, due to its oxygen storage properties [7]. Unfortunately, cerium oxide undergoes a fast surface area decrease when it is treated at high temperature [8]. Moreover, in several cases, this transition to lower surface area values occurs at temperatures that are lower than process operative conditions. Thus, in many cases, low surface area cerium-based catalysts need to be deposited onto structured supports [9]. The active phase usually consists of metal ions, and it is the core of the catalysis process. A variety of metals have been proposed in view of the different chemical processes. Among others, cobalt [10,11] and nickel [12,13] have been chosen because of their catalytic activity in many different

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reactions. Moreover, they have been proposed as a low cost and effective alternative to noble metals. A variety of geometrical supports are now commercially available for catalytic purposes. They mainly differ for the structural and physical properties, such as chemical nature, surface area per unit volume or mechanical properties. Three way catalysts (TWCs) are one of the most diffused and investigated applications of structured supports to catalysis [14] because they are extensively used for gas pollution control in vehicle exhaust [15]. Among the different supports, solid open-cell foams are highly promising. Open-cell foams are characterized by a high interface area and high porosity, which could result in lower pressure drops and higher energy efficiency [16,17,18]. Due to their geometrical structure, high performance in terms of fluid/solid mass transfer are guaranteed. As already reported, structured catalyst preparation is usually based on coating deposition onto the geometrical support surface. Depending on the structured substrate geometry, several deposition techniques for the catalytic thin layer have been made up [19,20]. Among them, dip-coating technique is the simplest, most versatile and cheapest one to be used in industrial practice. Moreover, dip-coating deposition can be easily applied to both metallic and ceramic supports of a variety of shapes [21]. The first step in dip-coating technique consists of the preparation of a slurry that is composed of a liquid phase in which the final powder to be deposited has to be suspended or better, dispersed. Then, the structured support is dipped in the liquid medium in order to fill the voids with a catalytic slurry precursor. Finally, the geometrical support is withdrawn from the slurry at a controlled rate. Depending on viscosity and support geometry, excess liquid removal is guaranteed by the opposite forces acting on the liquid during the withdrawal step. As a result, coating thickness depends on the balance between gravitational force, which promotes the removal of the liquid phase, and the viscous forces acting in the slurry, which involve the sliding resistance [22,23]. Therefore, the control of the coated layer properties, i.e., thickness and adhesion, are mainly ruled by the slurry rheological behavior and the withdrawal velocity. As far as rheological behavior is concerned, the suspension formulation (e.g., water/powder ratio, acid/powder ration, and surfactant content) can be easily tuned in order to achieve the desired viscosity at the typically applied dip coating shear rates [21]. Indeed, at low viscosity values, which promote good adhesion, lower coating loads are

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obtained. On the other hand, when a higher viscosity is applied, the higher coating load obtained is counterbalanced by poor adhesion, and sometimes a few difficulties are faced in applying the method [24]. Different methods are available for both powder suspension (or dispersion) and slurry stabilization. A well-known method is based on the generation of electrostatic repulsions among the powder particles, promoted by the surface charging of the material to be dispersed in an acidic environment [24,25,26]. This procedure has been applied with success to high surface area powders, as their surface can be easily charged simply by managing the pH of the suspension. Unfortunately, this route is not easily applicable in the case of powders with low surface area, those characterized by chemical unreactivity, or in the case of a possible dissolution of components. In all of these cases, an acid-free dispersion method needs to be used, and the selection of an alternative dispersant agent is required. Generally speaking, in the acidfree method organic molecules, more typically macromolecules, are used as dispersants. They are dissolved into the liquid phase so they can closely interact with the powder surface, allowing particle dispersion and slurry stabilization [27,28,29]. Many papers are reported in the literature regarding formulation and its effect on resulting slurries [30,31,32]. However, to our knowledge, fewer works have been reported on the production of slurries for coating deposition, via dipping, of low surface area powders onto open cell foams. In particular, scarce information is reported on the effect of the different dispersants and the composition of the final slurry properties. This study is of fundamental importance to correlate easily tunable experimental parameters, such as ratios between the components, with the rheological behavior of the slurry. Indeed, rheology is the main operative parameter to drive the final coating load, thickness, and adhesion, i.e., which are the parameters of interest for industrial application [33]. Accordingly, the aim of this work is to clarify these aspects by studying the formulations and the procedures to deposit thin active layers of low surface area model catalyst powder, such as Ni- or Co-supported, onto low surface area ZrO2. The catalytically active powders were produced by means of the incipient wetness impregnation technique using a commercial support. Different acid-free formulations based on water (H), glycerol (G), and polyvinyl alcohol (PVA) were studied, and the effect of the components that determine the slurry rheology was assessed. The different slurries were tested to coat 30 PPI (Pores Per Inch) ceramic open-cell foams via the dipcoating technique. A correlation among the final coating load, thickness, and adhesion after thermal treatments with the slurry composition and rheology

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was proposed. An attempt to rationalize the rheological behavior at the light of composition was also made.

RESULTS AND DISCUSSION Catalyst Characterization Catalytic powders characterization is reported in Figure 1 and Table 1. For the sake of comparison, the characterization of the pristine morphological carrier (ZrO2) is also reported.

XRD patterns (Figure 1) clearly showed the reflections of the monoclinic ZrO2 support. Additional reflections at 37.2°, 43.2° and 63° 2θ can be clearly seen in case of the Ni-containing sample; the latter were attributed to NiO, while the one that was detected at 36.9° in the Co-based sample was attributed to Co3O4.

Figure 1: XRD spectra of the impregnated powders. Table 1: Morphological characterization of the powders Sample

Active Phase Crystal Size (nm) (by XRD)

Surface Area (m2·g−1)

Pore Volume (cm3·g−1)

ZrLS

-

27

0.2

NiZrLS

28

14

0.1

CoZrLS

25

19

0.1

The crystallite dimensions of the carrier and the active phase, which were calculated according to Scherrer equation, were in the range of 25–28

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nm, thus, almost comparable. Regarding morphology, very close surface areas and pores volumes were measured for the active materials: a decrease of the surface area was observed upon impregnation that was accompanied by a decrease of the pore volumes. This effect is clearly explained by the partial occupation of pores due to the presence of the active phase. According to the procedure reported in the experimental section, all the powders, carrier and active materials were dispersed in the HGP liquid medium, and their rheological behaviors were compared. Results are reported in Figure 2.

Figure 2: Rheological behavior of zirconia-based slurries.

Apparently, active phase presence did not affect rheological properties. Indeed, shear thinning behaviors were found for all the samples. In order to get information on the non-Newtonianity degree of the slurries, the flow curves slope was determined in the 10–100 s−1 shear rate range. Generally speaking, higher slope values (in modulus) correspond to a more marked shear thinning behavior. For all samples, slopes in the range of 0.6–0.4 were calculated; this highlighted a very close rheological behavior for the three slurries. The powder composition exerted a different effect, since the presence of the active phase induced lower viscosity values in all the shear rate range (Figure 2). Such differences cannot be directly related neither to the surface area nor to the porosity due to the close values of these parameters detected for the three samples; similar considerations can be done for the active phase content, too. On the contrary, the absolute viscosity appeared to be much more specific for the metal cations present at the carrier surface. The effect of the presence of an active phase onto the carrier surface was already reported in the literature [34]. Accordingly, the modification of the

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surface nature was reported as responsible for a different powder-dispersant interaction, which influenced the rheological behavior. However, in this case, this effect should be quite peculiar considering the strong chemical similarity between Ni and Co both in terms of atomic weight and atomic number. Accordingly, an attempt to rationalize the viscosity behavior was done, considering properties other than those mentioned above or a combination of properties reported so far. A possible explanation for the rheological behavior could be found by considering the actual volume fraction of the powders. To evaluate this aspect, a representative model to describe the powders has to be built. In the model, three components were considered: (1) powder nature (mainly related to the molar weight); (2) powder density (mainly related to porosity); and (3) powder size and size distribution. The first consideration regarded the composition of the different powders, since slurries formulation was obtained by considering the powders content on mass base. The presence of the active phase in its oxidized form should be taken into account, since it affects the powder molar weight and consequently, the powder concentration in the slurry. In order to quantify this effect, the actual amount of the active phase (in the oxidized form (mOx. )) was determined according to Equation (1) Act.Ph

(1) where mMet.Act.Ph (g) is the mass of the active phase (metallic form), MMetalIon is the atomic mass of the metal ion and MOx the molecular weight of the active phase (oxidized form). Starting from Equation (1), support weight fraction was recalculated according to Equation (2). (2) K equals to one for the bare support, while it is lower than one for the active powders (Table 2). However, Kcannot completely describe the powders because no powder density was taken into account in Equation (1). Powder density is directly correlated to the morphological properties and particularly to porosity that can be evaluated by means of BET analysis. Therefore,

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in order to properly evaluate the actual density of the powders, crystallographic density and sample porosity were introduced in the calculation. In the case of active powders, crystallographic density was assumed as the average of the densities of all the phases, weighted by composition as reported in Equation (3):

(3) where “i” is the any phase present in the powder in the oxidized form. The “real” density was thus calculated by taking into account the evaluated material porosity. Porosity values were used to determine the powder void fraction (φ) and, thus, the real powder density (ρreal) calculated according to Equation (4): (4) Moreover, to give a complete description of the powders and to evaluate their actual concentration in the slurry, the particle size and size distribution, evaluated by granulometric analysis, were introduced in the model. Accordingly, the weighted volumetric fraction was calculated by Equation (5): (5) where Vfract is the volume of the particle (Vi) and Percentagei is the relative amount of particles with the i-th diameter. For volume calculation, a spherical shape was assumed for particles while the specific volume was assumed as equal to the reciprocal of density. Finally, the number of particles per unit volume was obtained by using Equation (6) (6) where ρreal is the real density of powder and Vliq is the volume of dispersing HGP liquid medium. Results are reported in Table 2. Table 2: Powder physical properties and slurry viscosity values Powder

K

Bulk Density (g·cm−3)

Real Density (g·cm−3)

Particles Concentration (Particles cm−3)

Viscosity (Pa·s) at Shear Rate: 1 s−1

10 s−1

100 s−1

ZrLS

1

5.68

1.13

1.50 × 108

0.743

0.185

0.067

NiZrLS

0.91

5.77

2.43

4.24 × 107

0.309

0.1

0.04

Washcoat Deposition of Ni- and Co-ZrO2 Low Surface Area Powders.... CoZrLS

0.76

5.78

2.16

4.10 × 107

0.532

0.14

139 0.059

On these bases, viscosity at three selected shear rates was plotted as a function of powder density (Figure 3). Viscosity was found to decrease linearly as powder density increased. Once the mass of powder to be dispersed is fixed, a larger material density results in a lower number of particles per volume unit. The number of particles per volume unit directly influences the rheological behavior: a higher number of particles leads to higher viscosity values [35,36]. Thus, the rheological slurry behavior can be explained on these bases. Pure ZrO2 (the carrier), which is a less dense material than the impregnated one, showed the highest viscosity because the same powder amount (weight base) had a larger number of particles present. When powder density increases, such as in the case of the active powders, the number of particles decreases; therefore, powder concentration and slurry viscosity also decreases.

Figure 3: Viscosity at three selected shear rates as the function of the powder density (triangles: SR 100 s−1; circles: SR 10 s−1; squares: SR 1 s−1).

This picture was confirmed by plotting the viscosity as a function of the particle concentration, which was calculated as reported above (Figure 4). At any shear rate, the viscosity increases with the particles concentration. Results of Figure 4 are in line with the conventional behavior of slurry viscosity. It is well known that the viscosity of concentrated dispersions is higher than that of diluted ones. This effect may be related with an increase

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of particle-particle interaction. Moreover, the fact that parallel lines were found suggested the presence of the same interaction mechanism at any shear rates.

Figure 4: Viscosity as a function of particles concentration (triangles: SR 100 s−1; circles: SR 10 s−1; squares: SR 1 s−1).

Washcoat Deposition HGP-based slurries were deposited onto the 30 PPI ceramic foams via dipcoating process and then, thermally treated. Up to three multiple depositions were performed.

Figure 5: Coating Load as a function of viscosity and of the number of subsequent depositions.

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It is well known that washcoat load, the parameter of interest, is correlated with viscosity, i.e., the main operative variable of the process [37]. Accordingly, the coating load after calcination was plotted as a function of the viscosity values at shear rate 10·s−1, the one of interest for dip-coating application (Figure 5). Regardless slurry formulation, a quite linear loadviscosity correlation was found. As expected, the washcoat load increased with viscosity and was observed at any dipping number. Due to the highest viscosity values, the higher loads were obtained with the pristine ZrO2 slurries which were found to reach approximately 22 wt. % after three dippings. When more than one dipping was applied, the deposition seemed not to be affected by the presence of the previous washcoat layer. This effect is better evidenced when the washcoat load after calcination is plotted as a function of the dipping number (Figure 6).

Figure 6: Coating load as a function of the dipping numbers.

A linear trend between load and dipping number was found, suggesting dipping number as a useful tool to manage washcoat load. Results on adhesion tests—by means of an ultrasound stress test in petroleum ether—are shown in Figure 7.

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Figure 7: Weight loss after adhesion test for Zr-based samples as a function of load after calcination.

From the analysis of points of Figure 7, adhesion seemed strongly related to the coating composition; in the case of pure ZrO2, negligible or no weight losses were present while the presence of active phase led to higher weight losses, with a relative maximum in correspondence of the intermediate washcoat load values. Generally speaking, higher washcoat load should determine higher losses after adhesion test, due to layer thickness. On the contrary, in these samples an overall increase of adhesion was found with load increase. For the considered samples, the effect of the different surface composition should also be taken into account even though this effect cannot be easily evaluated. Anyway, due to the large complexity of the involved phenomena and to the brittleness of the support foam, this point deserves more study to be clarified. A qualitative washcoat analysis as a function of dipping number was performed on the active powders by using an optical microscope (Figure 8). Figure 8 shows the images after flash drying. As described in the experimental section, this step is performed at 350 °C for 6 min: these operative parameters are suitable for solvent removal, but they are not enough for the total decomposition of the organic components which is still incomplete [38]. To analyze the washcoat at this step can be highly useful: the partial decomposition of the organic compound will result in a darkening of the surface that should allow for a better vision of coating coverage and homogeneity. Both samples clearly showed a surface darkening that was

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qualitatively interpreted as a homogeneous distribution of the washcoat load. As a matter of fact, white areas, which correspond to the bare substrate, were still evident only after one dipping; then, they gradually decreased and tended to disappear upon multiple depositions. After three dippings, a good coverage was reached. Almost no pore clogging occurred: only limited pore clogging was found for the CoZrLS (left side in the picture).

Figure 8: Optical images of NiZrLS (1D to3D) and CoZrLS (1D to3D) deposited on 30 PPI foams: effect of multiple dippings after flash drying. For the sake of comparison, an image of the bare foam has been added at the bottom of the image.

Figure 8 shows the images after flash drying. As described in the experimental section, this step is performed at 350 °C for 6 min: these operative parameters are suitable for solvent removal, but they are not enough for the total decomposition of the organic components which is still incomplete [38]. To analyze the washcoat at this step can be highly useful: the partial decomposition of the organic compound will result in a darkening of the surface that should allow for a better vision of coating coverage and homogeneity. Both samples clearly showed a surface darkening that was qualitatively interpreted as a homogeneous distribution of the washcoat load. As a matter of fact, white areas, which correspond to the bare substrate, were still evident only after one dipping; then, they gradually decreased and

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tended to disappear upon multiple depositions. After three dippings, a good coverage was reached. Almost no pore clogging occurred: only limited pore clogging was found for the CoZrLS (left side in the picture). In order to evaluate the deposited layers at higher magnification, washcoat after calcination was analyzed by SEM measurements (Figure 9). The results after three dippings are reported for the active powders. Acquisitions were performed in back scattering, and they once more demonstrated the good coverage homogeneity of the surface (Figure 9). Few defects were present, but they were of limited extent and localized. Coating surface, analyzed at higher magnification (Figure 9e,f), showed evidence of the presence of cracks of limited depth; another coating layer (not the bare support) was seen underneath.

Figure 9: Back scattering SEM analysis of ZrLS, (a,d), NiZrLS (b,e) and CoZrLS (c,f) coated samples after three dippings at different magnifications (100X and 1000X).

EXPERIMENTAL SECTION Catalytic Powders Preparation and Characterization A commercially available low surface area carrier was used for catalyst production, namely zirconium oxide supplied by Melcat (in the following, ZrLS). Catalysts were produced by using the incipient wetness impregnation method [19]. Nickel nitrate hexahydrate (98.5%, Sigma-Aldrich, St. Louis,

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MO, USA) and cobalt nitrate hexahydrate (98%, Sigma-Aldrich, St. Louis, MO, USA) were used as nickel and cobalt active phases precursors, respectively. These metal salts were used to obtain a final load equal to 7 wt. %, on a metal base. The precursors solution was wisely dropped onto the carrier. Impregnated powders were dried overnight at 120 °C, and then, they were calcinated in order to decompose the nitrate precursors and to obtain the final oxide. In particular, Co-based samples were calcinated at 400 °C while Ni-based samples were treated at 800 °C; in both cases, dwell was set at 10 h, with heating and cooling rate of 2 °C·min−1. The two different calcination temperatures were chosen in accordance with possible catalytic applications (i.e., Oxy-Steam Reforming for Ni-based catalysts [39] and Fisher Tropsch synthesis for Co-bases samples [11]. A final metal load about 10 wt. % was measured for all powders. Impregnated powders were characterized by means of X-ray diffraction. A D8 Advance diffractometer (Bruker, Billerica, MA, USA) and a CuKα radiation were used (10–80° 2θ range, 40 kV and 40 mA, step scan 0.02° 2θ, time 1 s·step−1). Crystallite dimensions were evaluated from the reflection line broadening (FWHM, calculated by Topas) using the Scherrer equation [40]. The powders particle size was evaluated by using a CILAS 1180 laser granulometer (Compagnie Industrielle des Lasers, Orleans, France). BET surface area and pore volume were determined by N2 adsorption and Hg intrusion; in the first case, a Tristar 3000 device was used (Micromeritics, Norcross, GA, USA). N2 physisorprion measurements were carried out after heating at 150 °C overnight, under vacuum. An Autopore IV instrument (Micromeritics, Norcross, GA, USA) was used for Hg intrusion.

Washcoating The powders dispersion formulation is based on a dispersant, glycerol (G) (87% w/w water solution, Sigma-Aldrich, St. Louis, MO, USA), a solvent/ diluent, distilled water (H), and a rheology modifier, polyvinyl alcohol (PVA) (Mowiol, Sigma-Aldrich, St. Louis, MO, USA). Weight ratios among the three components were calculated with respect to the powders (PW), and they were respectively set at: G/PW = 1.9, H/PW = 1.8 and PVA/PW = 0.07. In the following, the liquid medium based on this formulation will be labeled as HGP. The slurry was obtained by means of a procedure reported elsewhere [24]. Briefly, in a typical experiment, PVA was dissolved in distilled water at

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85 °C; then glycerol was added, always under magnetic stirring. The obtained HGP liquid medium was used to disperse catalyst powders. The powder was added to the HGP solution, and the resulting slurry was ball-milled for 24 h (50 rpm of rotation rate) in a polyethylene jar using ZrO2 spheres as grinding bodies. After the milling process, a sonication pre-treatment was performed for 30 min on the slurries in order to reduce foaming. The slurries rheological behavior was evaluated in the 1–103 s−1 shear rate range by means of a DSR 200 instrument (Rheometrics, New Castle, DE, USA) by using the parallel plates geometry and plates of 40 mm of diameter. Before coating deposition, supports were cleaned with acetone for 30 min in an ultrasound bath. Slurries were deposited on Yttria-stabilized Zirconia Alumina (YZA) open-cell foams (Selee Company, Hendersonville, NC, USA), with a nominal pore density of 30 PPI (pore per inch). Structured supports were cut in parallelepiped shape with squared section; dimensions were set at 1.5 cm and 1 cm for length and section, respectively. Dip-coating was used as the deposition technique. Both the dipping and withdrawal rate were set at 13 (cm·min−1). After the dipping step, coated samples were flash dried [38] for 6 min at 350 °C in a sealed oven. Then, a final calcination thermal treatment was performed for 10 h at 400 °C and 800 °C for Co- and Ni-based materials respectively; in both cases, dwell was set at 10 h with a heating and cooling rate of 2 °C·min−1. When necessary, multiple depositions were performed; the dipping procedure was repeated, and flash drying was performed between two subsequent dippings. Washcoat load was evaluated by the weight difference of bare and coated foam. Coated layers homogeneity and morphology were evaluated by means of optical (SZ-CTV microscope, Olympus, Tokio, Japan) and scanning electronic microscopy (Stereoscan 360, Cambridge Instruments microscope, Somerville, MA, USA). Coating adhesion was determined by coated samples sonication for 30 min in a petroleum ether bath, according to literature [34]. In Figure 10, a schematic representation of a typical procedure for structured catalyst production is reported.

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Figure 10: Scheme for catalyst production and washcoat deposition (slurry composition and details of the procedures as reported above).

CONCLUSIONS (1)

(2)

(3)

(4)

(5)

The use of an acid-free water-based formulation proved to be effective for the dispersion of model catalytic powders characterized by low surface areas. The obtained slurries are suitable for deposition on ceramic open cell foams via dipcoating. A dependence between the powder properties and the final rheology was evidenced. In particular, the rheological behavior, being directly related to both powder particles porosity and density, could be managed by taking into account these properties during formulation. Active phase presence influences particles properties, although present in a limited amount. From a practical point of view, chemical composition (i.e., molar weight) and porosity were found to be simple and detectable parameters to be used in slurry formulation to control the rheological behavior. The use of multiple dippings proved to fulfill the flexibility requirements for washcoat load management. All the formulations obtained good results in terms of washcoat load and adhesion. Local and limited pore clogging occurred only in the washcoat at very high load degree (higher than 20 wt. %). Depending on the powder to be deposited and on the operative condition, promising results have been obtained in terms of adhesion. Most samples displayed losses lower than 10 wt. %, which was reported in literature as satisfactory for washcoat

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(6)

adhesion on open-cell foams [38]. Although, up to now an “a priori” formulation of the slurry is hardly to be obtained without any experimental base, the results here reported can help to reduce experimental tests in a trial-anderror approach.

ACKNOWLEDGMENTS This work was funded by the Ministry of Education, University and Research, Italy (MIUR, Progetti di Ricerca Scientifica di Rilevante Interesse Nazionale 2010–2011) within the project IFOAMS (“Intensification of catalytic processes for clean energy, low-emission transport and sustainable chemistry using open cell FOAMS as novel advanced structured materials”, protocol no. 2010XFT2BB).

AUTHOR CONTRIBUTIONS R.B. and C.C. conceived and designed the experiments. M.C. and F.P. performed the experiments; R.B., M.C., F.P. and C.C. analyzed the data; C.C. contributed reagents/materials/analysis tools; R.B. and C.C. wrote the paper.

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SECTION IV: ROLL COATING AND RELATED PROCESSES

Chapter 10

Experiment Research on Hot-Rolling Processing of Nonsmooth Pit Surface

Yun-qing Gu1 , Tian-xing Fan1 , Jie-gang Mou1 , Wei-bo Yu2 , Gang Zhao3 , and Evan Wang4 College of Mechanical Engineering, Zhejiang University of Technology, Hangzhou 310014, China

1

China Aviation Powerplant Research Institute, Zhuzhou 412002, China

2

College of Mechanical and Electrical Engineering, Harbin Engineering University, Harbin 150001, China

3

Institut National Polytechnique of Grenoble, Joseph Fourier University, Grenoble 38031, France

4

ABSTRACT In order to achieve the nonsmooth surface drag reduction structure on the inner polymer coating of oil and gas pipelines and improve the efficiency

Citation: Yun-qing Gu, Tian-xing Fan, Jie-gang Mou, Wei-bo Yu, Gang Zhao, and Evan Wang, “Experiment Research on Hot-Rolling Processing of Nonsmooth Pit Surface,” Applied Bionics and Biomechanics, vol. 2016, Article ID 4915974, 10 pages, 2016. https:// doi.org/10.1155/2016/4915974 Copyright © 2016 Yun-qing Gu et al. This is an open access article distributed under the Creative Commons Attribution License, which permits unrestricted use, distribution, and reproduction in any medium, provided the original work is properly cited.

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of pipeline transport, a structural model of the machining robot on the pipe inner coating is established. Based on machining robot, an experimental technique is applied to research embossing and coating problems of rollinghead, and then the molding process rules under different conditions of rolling temperatures speeds and depth are analyzed. Also, an orthogonal experiment analysis method is employed to analyze the different effects of hot-rolling process apparatus on the embossed pits morphology and quality of rolling. The results also reveal that elevating the rolling temperature or decreasing the rolling speed can also improve the pit structure replication rates of the polymer coating surface, and the rolling feed has little effect on replication rates. After the rolling-head separates from the polymer coating, phenomenon of rebounding and refluxing of the polymer coating occurs, which is the reason of inability of the process. A continuous hot-rolling method for processing is used in the robot and the hot-rolling process of the processing apparatus is put in a dynamics analysis.

INTRODUCTION Continuous hot-rolling transfer process employs rollers for rolling and continuous imprinting. Its characteristics include better flatness, smaller imprint pressure, and requiring simpler equipment; moreover, it needs only one mold to complete large-area substrate continuously imprinting, which can increase the production greatly [1]. Hot-rolling transfer processes include two copy types, the first uses a cylindrical mold with fine surface characteristics, and the second type uses roller with smooth cylindrical surface. In the first type, fine surface characteristics can be achieved by using the traditional microfabrication on its surface or by the method of covering smooth cylindrical surface with flexible membrane. The second type employs the traditional flatbed mold to achieve continuous transfer process. Regarding the exploration of this process, it is still in its infancy [2, 3]. Hot-rolling transfer processes include three types, the platen, stepping, and continuous rolling. For some advantages like better flatness, smaller imprint pressure, simple equipment, and large-area imprinting ability over the other processes, continuous rolling processes are preferred in gravure and flexible flexographic printing technology [4] and they adopt this roll-to-roll (R2R) method, which is applied to manufacture flexible electrophoresis display and has already been tested for its higher reliability and process performance [5]. In the actual production process, the typical

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R2R method consists of three steps, deposition, transfer, and packaging [6]. Tan et al. [1] firstly proposed and demonstrated one process which applied roller continuous manufacturing method to transfer technology in micro/ nanomanufacturing area and presented the progress of roller nanoimprint lithography. Shen et al. [7] succeeded in producing 100~200 μm cylindrical microlens on the polycarbonate substrate by using the fine hot transfer process and studied the relationship between the lens size and imprinting pressure, temperature, and time during the microlens production process. Shan et al. [8, 9] successfully produced vertical microlenses and an optical conversion platform by using hot transfer technology. Their study showed that the embossed polycarbonate substrate temperature should be higher than 185°C, and imprinting pressure should be greater than 1.75 kg/mm2. Liang et al. [10] proposed a continuous process using roll-to-roll method. This continuous roll-to-roll rolling transfer process can manufacture products such as flexible displays at low cost and high efficiency. Chang et al. [11] proposed a method which can use air pressure to make silicon mold during the hot transfer process, when the air pressure, temperature, and pressure holding time was 10~40 kgf/cm2, 150°C, and 30~90 s, respectively. This could produce neat uniform and focusing powerful microlens array. Charest et al. [12] proposed a method of using hot transfer process to produce the morphology structure on the polymer substrate for cell growth and osteoblast growth direction response along the surface. In order to improve the release properties of Ni substrate, Byeon et al. [13] placed a film of SiO2 on the surface of Ni substrate and then coated a film of SAM (selfassembled monolayer) on the surface of SiO2. He found that it was very easy for the Ni substrate to be released during the polymer hot transfer. Air bubbles defects are often encountered during the manufacturing process of polymer. To avoid them, Mekaru et al. [14] added an ultrasonic vibration in longitudinal direction in the hot transfer process. This method greatly minimized the bubble defects. Lee et al. [15] incorporated a tiny capillary electrophoresis device on the surface of polymethyl methacrylate to separate the DNA sequences. The trench which was produced has good repeatability, whose error was less than 1%. Chang and Yang [16] proposed a method of using uniform pressure to load pressure, which can improve the accuracy of hot imprinting on the whole entire substrate. Wang et al. [17] developed a novel full-color electrophoretic membrane manufacturing process by using rolling method. An orthogonal test method was proposed by Li et al. [18] to search the relationship between hot transfer process conditions and transfer accuracy with imprinting temperature, pressure, and imprinting

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lead time. Experiment revealed that imprinting temperature and pressure determined the accuracy replication of fine groove depth; however, the accuracy replication of groove width and shape was determined by holding pressure and heating time. Lan et al. [19–21] considered the polymer material’s rheological properties changes with temperature and established a noncrystalline polymer material viscoelastic model near glass transition temperature and also built a simulation model of platen hot transfer process; moreover, they analyzed the effect rules between the process parameters and quality of transfer. Finally, for the large-area continuous transfer problems of fine characteristic structure, they proposed an evaluating method for continuous roller hot transfer quality using transfer replication rate and explored the effect rules between rolling process parameters and forming quality, which provided a guidance to the fine hot transfer process design of noncrystalline polymer substrate. In this paper, for the purpose of reducing resistance and energy of oil and gas pipelines, we present a manufacturing method for a nonsmooth surface reducing resistance structure on the surface of pipe inner-wall polymer coating. Also, we establish a pipe inner-wall machining robot model and apply the pipe inner-wall machining robot as carrier to analyze the problem of hot-rolling-head coating. In addition, we research the effects between different process parameters and hot-rolling results.

MODELING OF PIPE INNER-WALL MACHINING ROBOT The main job of modeling pipe inner-wall machining robot is to achieve micromachined structures drag reduction on the inner polymer coating surface of the oil and gas pipelines. The model has the capacity of walking, positioning, and high-precision microsurface processing inside the pipe. Modular design concept is used in the design model of pipe inner-wall machining robot in order to minimize the correlation between the various modules. According to its function, model of pipe inner-wall machining robot is divided into walking modules and processing modules. The walking modules are responsible for the entire robot to walk in the pipeline, and the processing module is responsible for roll processing the polymer coating on the surface of the pipe wall.

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Modeling of Walking Modules The structure of walking equipment is as follows [22]: (1) linear motors, (2) screw nut, (3) screw, (4) support module, (5) stepper motor, (6) adjusting screw, (7) walking wheel, and (8) adjusting nut. The required parameters of the walking equipment for pipe inner-wall machining robot are as follows:(1) powerful load capacity provides greater traction, (2) good stability, and (3) easy and simple control to achieve precise positioning. The equipment is peristaltic walkway, which can be divided into five parts, such as diameter adapting wheel mechanism 1, front support mechanism, peristaltic walk mechanism, back support mechanism, and diameter adapting wheel mechanism 2. In the walking equipment, diameter adapting mechanism 1 and front supporting mechanism constitute peristaltic front joint, diameter adapting mechanism 2 and back supporting mechanism constitute peristaltic back joint, and the equipment of peristaltic walk mechanism is in the middle of front and back joints. During the peristaltic walking, linear motors move to cause the elongation and contraction of the driving front and back joint, and each mechanism unites to achieve the robot’s walking function. The work steps of walking equipment are as follows. (1) Front support loosens followed by the clamping of the back support, and then linear motors elongation drives front joint to move forward. (2) Front support clamps, and back support loosens, and then linear motors contraction drives the back joint to move forward. (3) Repeat the previous two steps actions to achieve a peristaltic walking. Diameter adapting wheel mechanism is a hand tightening structure. Before machining the robot on the pipe inner-wall, manually adjusting the diameter wheel mechanism nut presses the nut against the pipe inner-wall. Front and back supporting mechanism is slider-crank mechanism under the control of screw nut. In the process of working, the driving motor drives the screw nut to make linear slide, and screw nut drives the crank to make swinging movement to achieve the support module’ pitching movement, which can achieve the support module clamping and releasing action in the pipe inner-wall. Diameter adapting wheel mechanism and supporting mechanism are both used symmetrically in the distributed architecture. It means that diameter adapting wheel mechanism has four groups for walking wheel structure with uniform distribution of 90° up, and supporting mechanism also has four groups to support modules with uniform distribution of 90° up. Diameter adapting wheel mechanism and crank slider support

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mechanism are both applied as truss structures, which possess powerful accommodating ability in pipe inner-wall, so in a certain range of pipe innerwall, this walking equipment can adjust its truss swing angle to adapt to those changes of pipe inner-wall.

Modeling of Processing Module The structure of processing equipment is as follows [22]: (1) DC servo motor, (2) driven bevel gear, (3) thrust ball bearing, (4) transmission shaft, (5) axle drive bevel pinion, (6) rolling tool, (7) heating rod, (8) slideway, (9) outer sleeve, (10) cam, (11) idler wheel, (12) cam, (13) outer sleeve, and (14) hollow shaft. Processing module is on the back of walking equipment and constitutes one part of walking equipment; the walking equipment can drag it to the processing position. This equipment includes three parts: front processing components, driving components, and back processing components. Driving components consist of DC servo motor and a series of bevel gears. Bevel gear system is a master-slave structure, which means a driving wheel drives two driven wheels reverse synchronous rotation. Front processing components and back processing components have the same symmetrical arrangement structure. The main parts of processing components are knife-head mechanism, slide track orienting mechanism, cam radial drive mechanism, and the inner and outer sleeve circumferential rotating mechanism. The distance of roll cutter is 190 mm between the two processing components. Processing components can complete two discontinuous processing on the polymer coating at one running. The length processed is 190 mm, which needs the back processing components to make two consecutive forward processing complete. In the processing of microdrag reduction surface structures, roll cutter rolled around the pipe axis suffers tangential reaction force, which passes to the processing component body at last, causing the body’s torsion deformation. It not only affects the structure’s stability, but also affects the pit’s direction processing and eventually leads to drag reduction structure which cannot achieve the desired drag reduction effect. In order to solve this problem, we use the method of dual tool counterrotating in the design. Transmission of bevel gear train can achieve synchronous reverse rotation of the left and right bevel gears connecting the cutter sets with the left and right bevel gears to achieve synchronous reverse rotation of two cutter sets. Under this processing mode, the force of tangential reaction force given to

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processing apparatus body can cancel each other and attains the equilibrium condition. During the processing, these two sets cutters rotate with left and right bevel gears, respectively, achieving the main movement. Under the role of cam acting and slide guiding, radial turret stretched along radial direction; it can achieve the cutters feed and retract movements in the processing. This structure enables the target single-power multifunction. It means one drive motor drives the entire process components to achieve the cutter’s main movement, feed movement, and retract movement, making the structure more compact with better action coordinating in the processing.

ANALYSIS OF HOT-ROLLING MACHINING ROBOT Hot-rolling-head is the main component of the pipe inner-wall machining robot, whose processing performance has a direct impact on the quality of the polymer coating surface.

Dynamics Analysis of Hot-Rolling Processing During the processing of the pipe inner-wall machining robot, the rolling cutter’s rolling movement on the pipe polymer coating surface can be simplified to an inner gear and an outer gear meshing process. The pipe polymer coating is equivalent to the inner gear, and rolling cutter is equivalent to the outer gear. The model of simplified processing module gear drive system is shown as in Figure 1. In a system of particles with ideal constraints, in any instant movement, all the active forces system, and inertial force system, the sum of the elementary work is zero in any virtual displacements; namely,

(1) If moment situation is considered, kinetic equation can be expressed as (2) From (1) and (2) we can get gear train transmission mechanism for kinetic equation as follows:

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Figure 1: The model of processing module gear drive system.

(3) where 𝐽1, 𝐽2, 𝐽3, 𝐽5, and 𝐽𝐻 are driving bevel gear 1, driven bevel gear 2, gear 3, gear 5, and 𝐻 carrier of inertia. 𝑚3 and 𝑚5 are the quality of gear 3 and gear 5. 1, 3, 5, and 𝐻 are the angular velocity of driven bevel gear 1, gear 3, gear 5, and carrier 𝐻. 1, 3, 5, and 𝐻 are the angular acceleration of driving bevel gear 1, gear 3, gear 5, and carrier 𝐻. 𝐹’ is the reaction force in the processing of roll cutter. 𝜂1 is the transmission efficiency between the outer sleeve rotating system from driving bevel gear and driven bevel gear. 𝜂2 is the transmission efficiency of roll cutter rotating mechanism. 𝑀 is the output torque of drive motor. From 𝑖3𝐻 = 3/ 𝐻 = 3/ 𝐻 = −𝑟/(𝑅 − 𝑟) the following can be obtained:

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(4) Gear 3’s structure and motion parameters are all the same as gear 5, so

(5) By putting (4), (5), and the following equation 𝑖1𝐻 = 𝑖12 = 1/ 𝐻 = 1/ 𝐻 𝑖 = −𝑧2/𝑧1 in (3), then we can get the kinetic equation of roll cutter (gear 3) as follows: (6) Equation (6) provides the gear train transmission dynamics expression of the pipe inner-wall machining robot processing module. This formula shows the relationship between angular acceleration 3 of roll cutter and the output torque 𝑀 of drive motor and reaction force 𝐹’ . Each component of the processing module is designed by considering the basics of the roll processing parameters, according to this relationship.

Analysis of Hot-Rolling-Head and Its Power The above-mentioned paragraph analyzes the overall structure of pipe innerwall machining robot. The main function of the machining robot is to apply a vector for hot-rolling-head, and the key core technology lies in how to roll out a nonsmooth surface structure with drag reduction effect, using hotrolling-head to press on the surface of pipe inner-wall polymer coating.

Surface Morphology of Rolling-Head The effect of drag reduction [23–26] on nonsmooth surface has already been recognized by domestic and foreign researchers; its main types are pit-type, scales-type, and triangular groove-type [27–31]. The structure size of pittype is two magnitudes bigger than triangular groove-type, and scales-type usually uses the method of biological replicates to produce. The production costs are high, and it is difficult to be applied on large scale [32, 33]. Considering the advantages of processing and maintenance and combining

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the actual situation of pipe wall machining robot, select the convex hull as the surface morphology of the rolling-head, and then achieve to roll out the pits characteristic structure on polymer coating surface. The rolling-head’s radius 𝑅 is 27 mm, and length 𝐿 is 90 mm. There is a through hole whose diameter is 36 mm in the inner rolling-head for placement of heating rods into it. Convex hull is hemispherical of radius 𝑟 = 0.8 mm and length ℎ = 0.8 mm, the convex hull is on the rolling-head and arranged in a diamond structure, and its axial spacing 𝑠 = 16 mm and circumferential angle 𝜃 = 20∘ . The model of hot-rolling-head is showed as in Figure 2.

Figure 2: Model of hot-rolling-head: (a) the morphology of rolling-head parameter and (b) the physical hot-rolling-head.

Power Analysis of Hot-Rolling-Head The heat requirement of the hot-rolling-head is start-up and holding temperature power. Start-up power 𝑊𝑞 is used for heating roller to change room temperature 𝑇1 to the required temperature 𝑇2. 𝑊𝑞 is composed of heating power and distribute power. 𝑞 is composed of heating power and distribute power. Heating power 𝑊1 is the power required to heat the hotrolling’s temperature from 𝑇1 to the required temperature 𝑇2, so formula 𝑊1 is (7)

where 𝐶 is the specific heat of hot rolling, 𝑀 is the quality of hot rolling, and 𝑡 is the heating time.

Distributing power is the power which hot-rolling lost during the temperature elevation because of conduction, radiation, and convection. Because the hot-

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rolling-head’s temperature always changes during heating time, it is difficult to compute distributed power theoretically. Therefore, the power of required temperature 𝑇2 multiplied by thermal coefficient 𝐾 constitutes the distribute power, as is shown in (8)

where 𝑊𝑐 is holding temperature power and 𝐾 is thermal coefficient.

Holding temperature power 𝑊𝑐 is the power to hold the hot-rolling-head’s temperature without changing, it is constituted by heat conduction power 𝑊𝑑, heat radiated power 𝑊𝑓, and heat convection power 𝑊 , and its formula is (9) The formula of heat conduction power 𝑊𝑑 is

(10)

where Δ𝑇 is the temperature difference between hot rollers and polymer coating, 𝐴 is the contact area between hot rollers and polymer coating, 𝜆 is the thermal conductivity, and 𝛿 is the roll head thickness. The formula of heat radiated power 𝑊𝑓 is

(11)

where 𝐴1 is the remaining hot-roller surface area which is contacted with the polymer coating, 𝜀 is the radiation coefficient of hot roller, and 𝑞ℎ is the heat radiation loss per unit area. The formula of heat convection power 𝑊𝑙 is

(12)

where 𝑎1 is convective heat loss dissipation factor and 𝜂 is convective heat loss correction factor. The formula of holding temperature power 𝑊𝑐 is

(13)

The formula of starting power is

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(14) Combined with the hot roller’s specific circumstances of pipe wall polymer coating, 𝑇1 = 18∘ C, 𝑇2 = 140∘ C, 𝐶 = 0.88 KJ/(kg⋅ ∘ C), 𝑀 = 0.41 kg, 𝐾 = 0.55∼0.65, 𝐴=3 cm2 , 𝐴1 = 136 cm2 , 𝑎1 = 0.124 W/cm2 , 𝜂=1, 𝜆 = 0.001 W/ (cm⋅ ∘ C), 𝑊𝑐 = 30 W, and 𝑊𝑞 = 210 W are determined. Considering the complexity during the actual processing on pipe wall polymer coating, the selected power should be greater than theoretical value. At last, the required heating power is 300 W. Based on the temperature sensor characteristics, select two angles to choose the temperature sensor. One is raising the temperature measurement accuracy, and another is reducing the temperature measurement error. Choosing thermocouple as temperature sensor, the measurement error of thermocouple is smaller than measurement error of thermal resistance. The hot-rolling-head’s temperature measurement range is 0~500°C and its display error does not exceed 0.5%. So relay output is selected, which could receive input thermocouple temperature sensor, and working environment is the oil and gas pipelines. Digital temperature control device is used.

TEXT SET AND TEST METHOD Hot-rolling crafts of pipe inner-wall machining robot are a continuous replication process, and its continuous replication process can also be applied to copy the characteristic structure of nonsmooth convex hull which has drag reduction effect on roller surface to the polymer coating surface in order to achieve a large-area nonsmooth surface replication. Based on the above theory, machining robot is built on the pipe inner-wall, and then the hot-rolling process is studied. Under different roll temperature, speed, and depth, the temperature and time dependence of the polymer material are analyzed.

Prototype Debugging Based on the above analysis, a test prototype of pipe inner-wall machining robot is built. After the assembly manufacturing of machining robot parts followed by prototype assembling, prototype debugging is needed to

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verify the control reliability of the motor and rationality of control strategy. Machining robot’s debugging work required the drive motor to achieve the required parameters. Machining robot walking modules have 57H56 DC stepper motor and TG-300-D motorized faders, machining robot processing module has 90BLDC servo motor, and the debugging contains hardware, software, and system debugging. Hardware debugging contains power test, I/O connection check, and safety circuit confirmation. For programming system, we use software STEP7. First, we achieve the interface between the computer and the PLC and ensure proper communication between the two. Then, we download the system to the PLC and open the actual operation of the online monitoring. At last, we compared the PLC actual data with the project data and made proper adjustments. System debugging is the whole debugging between control systems and pipe inner-wall machining robot. During the debugging, the rule is to meet the control requirements and ensure system security. In order to select the optimal control parameters, first, we let the robot walk in the pipeline, and then, according to the pipeline diameter and walking displacement, we adjust the pulse volume and analog observation.

Coating of Polymer Pipe Inner-Wall The main job of inner-wall pipelines coating is to protect oil and gas pipelines against transporting media’s corrosion and to reduce the resistance of transporting media. Gas medium is a combination of a mixture of combustible and noncombustible material, whose main component is a substance of low molecular weight saturated with hydrocarbons, and some small amounts of nonhydrocarbon substances [34, 35]. Since some acidic substance constitutes the medium, and it can corrode the inner-wall, a polyurethane category test is selected at room temperature for curable coating taking requirements of oil and gas pipelines drag-reducing into account. This coating is composed of polyisocyanate prepolymer, fillers, additives, solvents, and some other materials. And it has good wear resistance and good adhesion, with good water resistance, moisture resistance, oil resistance, good acid and alkali resistance, solvent resistance, and chemical resistance, so it is very good for the drag reduction and corrosion on pipe coating equipment. The process of polymer coating is as follows:(1)Cleaning of pipe innerwall: the dirt and other attachments are wiped on the pipe inner-wall by using mechanical cleaning methods, solvents or high-pressure water, and so forth. Then the pipeline is preheated to temperature 3°C higher than the

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room temperature, and shot blasting is applied to do the rust treatment to achieve the Sa2.5 class. Finally compressed air is employed for purging the sand and dust off the pipe inner-wall.(2)Coating of pipe inner-wall: the inner-wall is coated manually in order to maintain the coating thickness uniformity; multiple coatings are desirable with a brush whose thickness is determined by testing requirements. Finally, the paint is sprayed to the pipe inner-wall by an airless spray pump.(3)The handle of coating solidification: the coating solidification usually relies on solvent evaporation or coating chemical reaction in terms of the actual situation of pipeline, but we use the hot air to make coating solidification faster.

Design Conditions of Hot-Rolling Process The hot-rolling process is widely used in the field of micromanufacturing as a manufacturing technique. However, in order to achieve large-area nonsmooth surface replication molding technology as a continuous rolling, its process control parameters are still in the research and practice phases. The main affecting factors of hot-rolling process parameters are temperature, time, speed, and pressure. Since the rolling-head of machining robot uses the cam’s lift and return to achieve the rolling-head for carrying out the plate polymer coating, different temperature, rolling speed, rolling time, and rolling are mainly considered for analysis of the effect of rolling. Hot-rolling process mainly includes three phases: the preheating, hotrolling, and cooling. The temperature curve versus time in the process is showed as in Figure 3. Specific steps are as follows: in the phases of preheating, the rollinghead’s temperature is raised to the temperature of polymer glass transition temperature above 𝑇1, then maintaining this temperature for a while 𝑆𝑇1.The polymer’s state is viscoelastic state at this temperature. In the phases of hot rolling, to roll the pipeline polymer coating by cam’s lifting for a rolling depth ℎ and keep the rolling time for a while ℎ𝑇1, the pressure of coating surface becomes smaller as the temperature of the polymer coating decreases at this time, and the temperature elevates through the conduction heating between hot rollers and coating. High temperature is maintained in some localized parts of the coating surface only, to ensure a faster mobility, and the bottom’s temperature is low enough, so there is no mobility produced on the bottom of coating. In the hot cooling phase, the temperature of rolling-head and polymer pipe inner-wall is lowered to the glass transition temperature below 𝑇2 instantaneously. Finally, the coating is cooled in the natural environment, and its temperature slowly returned to normal; in addition, rolling-head separates from the polymer coating on the

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cam’s return.

Figure 3: Curve of temperature versus that in hot-rolling process.

Replication Rate and Measurement In order to evaluate replication results under different hot-rolling process parameters, we define the replication rate as the final pits structure characteristics of the coating surface, which means that the replication rate is the volume ratio between all pit features and standard pit features after the forming rolling. It is very difficult to accurately measure and calculate pits structure characteristics, so the laser displacement scanner (type FT5070F) is used to measure sectional shape of pits structure characteristics on polymer coating. This can change the replication rate to area ratio between the intermediate sectional area of pits structure characteristics on polymer coating and the intermediate sectional area of standard pits structure characteristic: (15) where S’ is the intermediate sectional area of pits structure characteristics on polymer coating and S is the intermediate sectional area of standard pits structure characteristic. The detection distance of laser displacement scanner is 30~100 mm, and its accuracy is 0.03~0.1 mm (0.1% of MBE). Light source

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is red laser 650 nm, and working temperature is −10~60°C. This scanner is connected to the computer through the serial port and got the data acquisition by serial debugging software, and this scanner calculated the replication rate by measuring the sectional shape of rolling pits structure characteristics in processing. The distance between test pits and reference points is measured by laser displacement scanner and the output is communicated with the computer. Also, the measured data is processed by MATLAB software, and it restores the sectional shape of rolling characteristic structure and then calculates the sectional area by the method of trapz function.

TEST RESULTS AND ANALYSIS Control parameters in the hot-rolling process are as follows: the rolling temperature is 150°C, rolling depth is 0.8 mm, and rolling speed is 1 rad/s. The formed pits sectional morphologies are shown in Figure 4. It compares the characteristic structure of hot-rolling-head with the pits characteristic structure of polymer coating. It can be inferred that pits structure replication can be achieved on polymer coating, and consistent lines morphology can be achieved with the standard pits. Through formula (15), we calculate that the replication rate is 85.0%, which means that the replication rate is higher enough, and it also can verify the design rationality and the processing accuracy of the pipe inner-wall machining robot.

Figure 4: The comparison chart between actual intermediate sectional area and ideal intermediate sectional area: (1) ideal pit morphology and (2) actual pit morphology.

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In order to study the effect of the rolling process parameters on replication, different rolling depth ℎ, rolling speed 𝜔, and rolling temperature 𝑇 are chosen to have a series of tests. Under the condition of ℎ = 0.2 mm, 𝑇 = 140∘ C, 145∘ C, 150∘ C, 𝜔 = 0.5 rad/s, and 1 rad/s, 1.5 rad/s, we start a rolling test and measure and calculate the hot-rolling replication rate. After that, we develop a variable chart for varying replication rate and rolling temperature, as is shown in Figure 5.

Figure 5: The changing chart in which replication rate changed with rolling temperature.

The histogram in Figure 5 shows that under the same rolling depth, replication rate increases at elevated temperature, and at constant temperature, replication rate increases at lowered rolling speed. The polymer is very sensitive to temperature changing, so its mechanical properties is viscoelastic above the glass transition temperature, and its stress relaxation speed also increases, as its creep speed increases with increasing temperature. When the rolling-head is processed on the polymer coating, elastic deformation occurred firstly, followed by viscous deformation; viscous deformation is much bigger than elastic deformation. At the end time of 𝑇 = 140°C, 145°C, and 150°C, there is a proper amount of rebound phenomena of the pit morphology on polymer coating, due to the elastic deformation when hot-rolling-head separates from polymer coating. When the polymer

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coating temperature is lowered below the glass transition temperature, its mechanical state is glassy, and the rebound phenomenon stops. So the accuracy of replication can be improved by increasing temperature or decreasing rolling speed in the actual process. Under the circumstance of 𝑇 = 150∘ C, ℎ = 0.2 mm, 0.4 mm, 0.6 mm, 𝜔 = 0.5 rad/s, 1 rad/s, and 1.5 rad/s another rolling test is conducted and then measured and calculated the hot-rolling replication rate. We develop another chart in which replication rate changed with rolling depth, as is showed in Figure 6. As it can be seen from the histogram in Figure 6, replication rate increases with decreasing rolling speed, which shows good time-dependency of polymer coating, under the same rolling depth. Replication rate increases with increasing rolling depth at low rolling speed, with obvious change of replication rate. At higher rolling speed, replication rate increases slightly with the changing rolling depth. It shows that increase of the rolling depth cannot increase the replication accuracy significantly.

Figure 6: A changing chart in which replication rate changed with rolling depth.

In order to elaborate further the effect of the three main control parameters on replication rate in hot-rolling process, the orthogonal method is used to analyze the parameters in hot-rolling process. Factor A is rolling depth, factor B is rolling speed, and factor C is rolling temperature, as shown in Table 1.

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Table 1: Three factors and 3 horizontal test tables

In Table 1, three factors and three horizontal test tables are created with a total of 33 = 27 times according to 𝐿9(34 ), selected 9 from 27, and those selected numbers are 1, 5, 9, 11, 15, 16, 21, 22, and 26, and then test those selected numbers under new ID numbers 1 to 9. In experimental tests, 1/3 of total test numbers are done. In order to ensure the stability of the test results, the tests are repeated three times in each case, and then we get three factors coupling test results 𝐴, 𝐵, and 𝐶, as shown in Table 2. In Table 2, 𝐻1 represents the total average values of replication rate in order factors 𝐴, 𝐵, and 𝐶 in the first level of the test; 𝐻2 represents the total average values of replication rate in order factors 𝐴, 𝐵, and𝐶in the second level of the test; 𝐻3 represents the total average values of replication rate in order factors 𝐴, 𝐵, and 𝐶 in the third level of the test. So ℎ1 = 𝐻1/3, ℎ2 = 𝐻2/3, and ℎ3 = 𝐻3/3 are the average value of each corresponding level. In the same column, the maximum value in ℎ1, ℎ2, ℎ3 subtracts the minimum value which equals the 𝑋max − 𝑋min. By changing the factors listed in the maximum level table, the test target replication rate also acquires maximum change and, conversely, makes minimum change. Table 2 shows that factor 𝐴 has the smallest different value (1.1%), so the level of factor 𝐴 changes, which has minimal impact on the replication rate. The third level of factor 𝐴 corresponds to the maximum average value of replication rate as 62.1%, so the third level of factor 𝐴 is the best. Factor 𝐵 has the biggest different value (34.8%), so if the level of factor 𝐵 changes, it has maximum impact on the replication rate; meanwhile, factor 𝐵 should be used as a major factor in the study.

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Table 2: The coupling test analysis table of three factors A, B, C and

The third level of factor 𝐵 corresponding to the maximum average value of replication rate is 78.7%, so the third level of factor 𝐵 is the best. The different value of factor 𝐶 is 10.7%, and the third level of factor 𝐶 corresponds to the maximum average value of replication rate as 67.4%, so the third level of factor 𝐶 is the best. So the three factors effects on the replication rate are ranked as 𝐵, 𝐶, 𝐴, and the best program is number 3 in Table 2, whose replication rate is 83.8%, in which 𝐵3: rolling speed, the third level, and 0.5 rad/s, 𝐶3: rolling temperature, the third level and 150∘ C, and 𝐴3: rolling depth, the third level, and 0.6 mm.

As the best program 𝐵3𝐶3𝐴3 is not listed in the total 9 tests, the program 𝐵3𝐶3𝐴3 is tested to verify whether we can get the best results by orthogonal experiment method, then the calculated replication rate is 87.9% by formula (15), and the replication rate is higher than what replication rate obtains in test number 3, which shows that the best program 𝐵3𝐶3𝐴3 by orthogonal experiment complied with the actual situation. It has been found through the above analysis that the best replication rate is 87.9% and realized the pit-type

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replication of nonsmooth surface. The reason that the hot-roller processing cannot achieve complete replication of pit-type nonsmooth surface is that when hot-roller head is in contact with polymer coating, its temperature is higher than temperature Tg in which the glass state transits to viscoelastic state, and polymer coating is in viscoelastic state. When the two separates, polymer coating will be cooling with natural surrounding environment temperature, and polymeric materials will appear the rebound phenomena, which leads pit structure to tiny deformation, changing pit morphology, and then reducing the replication rate.

CONCLUSIONS The structural model of walking modules and processing module for the pipe inner-wall machining robot is established and their work process is analyzed. What is more, kinetic equations of hot-rolling process are also established, and the relationships between angular velocity, output torque, and rolling reaction force of rolling cutter are obtained. In addition, surface morphology of the roller-head is designed, and the principle of selecting roller-head power is given. The experimental prototype of pipe inner-wall machining robot is built, and the method of coating the polymer pipe inner-wall is given, and then the conditions of hot-rolling processing are designed. Experimental method is used to analyze hot-rolling effects under different process parameters. Meanwhile, it is also verified that the rolling temperature will be elevated and the speed of rolling will be decreased. It shows that both can improve the pit structure replication rates of the polymer coating surface, and the rolling depth has little effect on replication rates. The effect of hot-rolling process parameters is analyzed on pit morphology through orthogonal method. According to effect of replication rate on various factors, the ranked order is obtained: rolling speed, rolling temperature, and rolling depth. The best level program is B3C3A3 with replication rate of 87.9%. The reason that hot-roller processing cannot achieve the copy of the pit-type nonsmooth surface completely is the viscoelastic state of polymer coating. When hot-rolling-head separates from polymer coating, polymer coating encountered rebound phenomena, which leads pit structure to tiny deformation and reduces the replication rate.

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ACKNOWLEDGMENT The work was supported by the Zhejiang Provincial Natural Science Foundation of China (Grant no. LQ15E050005).

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Chapter 11

Bioinspired Superhydrophobic Surfaces, Fabricated Through Simple and Scalable Roll-To-Roll Processing

Sung-Hoon Park1, Sangeui Lee2, David Moreira3 , Prabhakar R. Bandaru3, InTaek Han2 & Dong-JinYun2 Department of Mechanical engineering, Soongsil University, 369 Sangdo-ro, Dongjak-gu, Seoul, 156-743, Korea

1

Material Research Center, Samsung Advanced Institute of Technology, Yongin-si, Gyeonggi-do, 446-712, Korea

2

Department of Mechanical & Aerospace Engineering, University of California, San Diego, La Jolla, CA 92093- 0411, USA 3

ABSTRACT A simple, scalable, non-lithographic, technique for fabricating durable superhydrophobic (SH) surfaces, based on the fingering instabilities Citation: Park, S.-H. et al. Bioinspired superhydrophobic surfaces, fabricated through simple and scalable roll-to-roll processing. Sci. Rep.5, 15430 (2015). https://doi.org/10.1038/ srep15430 Copyright © The Author(s) 2015. This work is licensed under a Creative Commons Attribution 4.0 International License. The images or other third party material in this article are included in the article’s Creative Commons license, unless indicated otherwise in the credit line; if the material is not included under the Creative Commons license, users will need to obtain permission from the license holder to reproduce the material. To view a copy of this license, visit http://creativecommons.org/licenses/by/4.0/

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associated with non-Newtonian flow and shear tearing, has been developed. The high viscosity of the nanotube/elastomer paste has been exploited for the fabrication. The fabricated SH surfaces had the appearance of bristled shark skin and were robust with respect to mechanical forces. While flow instability is regarded as adverse to roll-coating processes for fabricating uniform films, we especially use the effect to create the SH surface. Along with their durability and self-cleaning capabilities, we have demonstrated drag reduction effects of the fabricated films through dynamic flow measurements.

INTRODUCTION Superhydrophobic (SH) surfaces, i.e. surfaces with a water contact angle (WCA) above 150° and a hysteresis angle lower than 5°, have are of extensive interest in various scientific and engineering fields1,2, with various applications including self-cleaning surfaces, anti-icing coatings, antiadhesion coatings, and microfluidic systems3,4,5,6,7,8. Such surfaces are typically created by coating substrates with low-surface–energy materials coupled with controlling the surface roughness at both the micro- and nanoscales9. While a variety of approaches to SH behavior have been developed10,11,12,13,14, there are still critical barriers to their widespread use due to the cost, number of processing steps, limits on the manufacturable area, durability etc. For example, a soft-lithographic imprinting method using a polydimethylsiloxane (PDMS) stamp, which was prepared by replica molding against a hydrophobic lotus-leaf surface, was developed for SH surfaces15. Using such processes, the highest WCAs were achieved when patterns with high aspect ratio were used. However, such patterns were easily damaged through tribological interactions. Recently, designs incorporating a dual-hole pattern and a hierarchical micro-/nano-structure were developed for robustness against tribological damage16,17. However, such methods involved require a complicated series of steps and are consequently limited to small area substrates. With proper processing temperatures, solventcasting methods could also be used to form SH coatings on various substrates18. Furthermore, in conjunction with conjugated conducting

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polymers, such methods could also be used to create electrically active SH coatings. However, durability and uniformity yet remain key issues restricting commercial application. In other approaches, nanoparticle/fiber coatings on textiles have produced robust SH surfaces; this is possibly the method most amenable to practical applications19,20,21, but such methods are not applicable in all situations due to the requirement that the fabric have microscale roughness22. Additionally, laminate-templating methods using a membrane or mesh plate were developed to produce SH surfaces with abrasion resistance by laminating thermoplastic film and microscale mesh template23,24,25, heating under pressure to near the film’s melting temperature, cooling, and then peeling off the template. Hair-like SH surface structures were developed through the capillary effects between the mesh and film. However, such methods are limited yet by the heating cycles, the production time and issues related to the template reusability. Through a careful consideration of such state-of-the art techniques, we developed a simple as well as scalable roll-coating based process for fabricating SH surfaces harnessing non-Newtonian flow instabilities and associated effects, such as ribbing26,27,28,29,30. Our method has the advantage of requiring no additional chemical, vacuum, heat-treatment, or cleaning steps, nor any patterned template such as a master stamp or mesh. In addition to the efficiency and scalability of the process, the resulting tribologically stable SH surface has a bristled shark-skin like structure which could facilitate drag reduction effects in dynamic liquid flow over the surface.

RESULTS Formation of the Bristled Shark-skin like SH Pattern A double-roll based process was used to fabricate the SH film, as schematically shown in Fig. 1a. A high viscosity paste (comprised of 10 wt% multiwalled carbon nanotubes (MWCNTs) dispersed into polydimethylsiloxane (PDMS)) was placed in between the rollers, and the paste was transferred onto the roll with the higher speed, e.g., roll #1 (rotating at a velocity of V1) and roll #2 (rotating at a velocity of V2). Where if V1 > V2, the paste was coated as a smooth film (with less than 300nm roughness) onto roll #1.

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Figure 1: Fabrication of the shark-skin–like patterned films. (a) Schematic diagram of the double-roll film-making machine. (b) Pattern formation results when capillary bridges form between the rolls when their rotation speeds are inverted. (c) SEM images of the pattern produced by the rolling process. In inset image, individual MWCNTs were observed at the end of the pattern tip.

When V2 exceeds V1, it was observed that the paste on roll #1 starts to stick and then transfer to roll #2 in the nip area between the two rolls by capillary bridging31, as shown Fig. 1b. At a certain V2/V1 ratio (>1), a randomly structured pattern develops over the entire film area as indicated in Fig. 1c. Additionally, individual MWCNTs were observed at the end of the pattern tip due to capillary bridge effects. The resulting pattern, consisting of micro- and nano-scale features, resembles that of bristled shark skin with comparable scales and suggests possible use for a low drag material32,33. Demonstrating the large-scale uniformity of the films. High shear/yield stress (τ0) and thixotropic viscosities, important for maintaining the pattern structure, were measured (10 wt% MWCNT paste), as indicated in Fig. 2a,b26,34,35,36. For the fabrication of high-viscosity paste, three-roll milling techniques were used to disentangle and disperse high aspect ratio MWCNTs in the polymer matrix37. The paste was under shear force in the nip region between the roller contact area (Fig. 1b). Then, the paste was not sheared after rolling out from the nip between the rollers, which means the shear rate was nearly zero, and the developed morphology was unchanged by the yield stress. The non-Newtonian shear stress and

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viscosity of the MWCNT paste were shown to obey the Herschel-Bulkley equation35. (1) where, τ0 is the yield stress, η is the shear rate (1/sec), and K and n are constants. The value of n was less than 1.0, which implies the MWCNT/ PDMS paste has a shear-thinning behavior, shown in Fig. 2b.

Figure 2: Properties of the high viscosity MWCNT paste and the developed patterns. (a) Shear stress and (b) shear viscosity of the 10 wt% MWCNT/PDMS paste at various shear rates. (c) WCA and SA properties of the patterns generated at different roll shear rates (η). (d) Roughness of the patterns generated at different roll shear rates.

For a given film thickness and viscosity, we could optimize the pattern of the SH surface. The surface topology could be controlled through the effective shear rate η = ΔV/h between the two rolls, where ΔV = V1 − V2and h is the thickness of film. A smooth film without any patterning was formed when η > 0 s−1 (e.g., for η = 32 s−1, V1 > V2), as shown in Fig. 3a, and irregular surface morphologies were observed when η < 0 s−1, as shown in Fig. 3b–d. At a certain shear rate (e.g., for η = −82 s−1, V1 < V2), a uniform SH shark-

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skin–like pattern developed, as shown in Fig. 3c,g. The pattern morphology possesses larger features beyond the optimal shear rate (η > −82 s−1) for SH surface, resulting in a loss of hydrophobicity, as shown in Fig. 3d,h.

Figure 3: Surface morphology and contact angle image of the patterns produced at different shear rates. Tilted SEM images of films fabricated at (a) η = 32 s–1, (b) η = −44 s–1, (c) η = −82 s–1, and (d) η = −106 s–1. The corresponding WCA behaviors are in (e–h).

Figure 2c showed the relationship between the shear rate between the two rolls and the WCA. When V2 is larger than V1, WCA is gradually increased up to 161.3° at a shear rate of η = −82 s−1, while hydrophobicity is decreased after η < −82 s−1. The sliding angle (SA) is another important parameter describing SH surfaces and can be measured as the angle where the droplet rolls off when tilting the substrate. With similar trend of WCA, minimum SA (below 5°) is observed at a shear rate of η = −82 s−1indicating a selfcleaning surface as shown Fig. 2c. In addition, we measured the advancing angle, 155°, and the receding angle, 150° at a shear rate of η = −82 s−1 by dynamic sessile drop method. The hysteresis was around 5°. Lager negative η values (i.e., V1 < V2) increased the surface roughness, as shown in Fig. 2d. At even higher shear rates (i.e., with V1 ≪ V2), the entire film was transferred from roll #1 to roll #2 observing starvation of the paste on roll #1.

Mechanism underlying the Formation of a Bioinspired SH Surface Instability in Newtonian or non-Newtonian flow was typically regarded as having a negative impact on roll-coating processes for fabricating uniform films due to ribbing or zigzaging26,27,28,29,30. However, in our work we

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have exploited such characteristics to pattern the SH surface. A meniscus forms in the fluid region where two neighboring rollers separate, and the fluid runs on the rollers as shown in Fig. 1b. Instability in the meniscus occurs when the shear force and surface tension are not balanced. While the fingering instability of Newtonian fluids has been investigated analytically and experimentally26,27,28,29,38,39, there have been few studies on the related instabilities for non-Newtonian flows, for liquids with extremely high viscosity, as for the paste used in our study. An apparent morphological change from flat to rough pattern (Fig. 3a,b) was observed at a shear rate between 32 s−1 and −44 s−1. It is claimed that the creation of rough surface pattern is directly related to the sign change in V1 − V2, that is, V1 − V2 > 0 in Fig. 3a, whereas V1 − V2 < 0 in Fig. 3b–d. According to Grillet’s study28, treelike structures were observed in eccentric cylinder forward roll coating flow for both Newtonian and non-Newtonian fluids, while not in the eccentric roll-and-plate flow. The speed of the outside roll (say V2) is higher than that of the inside roll (say V1) in the eccentric forward roll coating (V1 − V2 < 0), while the speed of the outside roll is less than that of the inside roll in the roll-and-plate coating (V1 − V2 > 0).

Figure 4: Formation mechanism of the shark-skin–like pattern. (a) Schematic diagram showing the relationship between the pressure gradient at the advancing front and surface energy26. (b) Schematic of the velocity and pressure gradient component radiating outward in the z-direction, tip splitting, and tip

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coalescence. (c) Top-view SEM images of the patterned surface fabricated at η = −82 s−1. Inset indicates low resolution image.

We indicate that fingering instabilities associated with non-Newtonian flows between the rollers are necessary for the formation of the observed rough patterns as shown in Fig. 426,27,28,29,38,39. The flow velocity (v0) is proportional to the pressure gradient (dp/dx)0 at the meniscus front, and the curvature of the front (R) and the surface energy (γ) determine the flow-surface geometry, as in Fig. 4a. The R = λ2/(4π2a) in the case where the meniscus is perturbed into a sinusoidal shape with amplitude a and wavelength λ. If the unperturbed velocity profile is assumed to be v0 = − (h2/12μ)·(dp/dx)026, where μ is the shear viscosity of the paste, the front instability will occur under conditions where −a(dp/dx)0 > γ /R. With:

(2) where Ca = |μv0/γ| is the capillary number and v0 = (V1 + V2)/2. The Cadescribes the relative influences of the viscous and surface forces in the instability region, where the former causes the instability, while the latter acts as a stabilizer. The velocity of each roll, the effective shear rate and Ca produced between the two rolls are summarized. Under optimal conditions where WCA reaches the maximum in Fig. 3c, Ca was very high26,27,28,29,38,39 at ~ 2700, from the values in Fig. 2b (μ = 658 Pa·s). Here, v0 = 4.1 cm/s, where v0was estimated from the speeds of the two rollers, (V1 + V2)/2 (=Vavg), and a γ- value of 21.8 mNm−1 for uncured MWCNT paste film, as determined using the Owens-Wendt method40. Such a high Ca in our study was primarily attributed to the high viscosity of MWCNT paste. The measured microstructure sizes (λ) reached around 60μm from top view of the pattern images in Fig. 4c . The average height of the SH structures was 30 ~ 40 μm. The calculated microstructure sizes (λ) from equation (2) was 17.4 μm, close to the size observed in the actual surfaces. In addition, the expectation from the equation (2) that the increase in the pattern size with decreasing Ca was consistent with our observation as shown in Fig. 3. The relative error between the observation and the calculation of λ from Equation 2 was believed to result from the assumption that the velocity was established under incompressible and Newtonian flow conditions. There may be significant differences between the velocity profile modeled for Equation

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(2) and that of the non-Newtonian MWCNT/PDMS paste by shear-thinning phenomenon35. In addition, velocity and pressure components in spanwise direction of the rolls in Fig. 4a,b increase as the capillary number (or the viscosity) increases26,28,29. As the flow velocity or the capillary parameter increase, the fingering structures grow parallel to circumferential direction of the roller and eventually to neighboring fingers, and then form branched or tree-like structures26,28,29. The tip-splitting can be treated as signature of the velocity components radiating outward to the spanwise direction of the rollers due to the pressure gradients as shown in Fig. 4b,c. However, the velocity profile has been assumed to have a component in the circumferential direction only.

Durability Performance of the SH Films To evaluate wear durability for practical usage, a rubber tip under an applied normal load (of ~1.5 N) was dragged horizontally on the fabricated SH surface and on a pillar-type patterned surface. As shown in Fig. 5, the latter patterns were found to be severely damaged after 100 cycles, resulting in a considerable degradation of hydrophobicity (inset of Fig. 5). The produced rough/shark-skin–like pattern, on the other hand, maintained its morphology and hydrophobicity even after 2000 cycles, demonstrating its robustness to tribological/physical damage16,25,41.

Figure 5: Mechanical durability of the SH surfaces. WCA values of the sharkskin–like pattern and a pillar-type pattern after N wear-test cycles by a rubber tip under an applied normal load of 1.5 N. The pillar-type surfaces, shown inset, are severely damaged after 100 cycles and show significant degradation of hydrophobicity. The shark-skin–like surface maintains its contact angle and morphology after 2000 cycles.

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Drag Reduction Performance of the SH Film Generally, the relative velocity at the boundary between a solid wall and a liquid is considered to be zero, through the widely accepted “no-slip” boundary condition. However, for the case of SH surfaces, the fluid velocity near the solid surface appears to exhibit a non-zero velocity when averaged over the complete surface, taking into account the air/fluid areas, i.e., induced through a layer of air that can form in the interstices of the rough surface (inset scheme of Fig. 6), which acts as a shear-reducing boundary between the solid surface and the fluid, greatly reducing the drag42,43.

Figure 6: (a) Water velocity profiles for water flow on SH patterned surfaces and (b) Pressure drop through a channel in accordance with the velocity measurements. The lower surface is SH while the upper surface is a glass coverslip. Data points (#, •) correspond to the fluid velocity averaged over the field of view at each height during two experimental runs under the same conditions.

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The solid line corresponds to a fitted pressure-driven Poiseuille flow profile with a slip boundary condition applied to the lower surface. Upper inset shows silvery mirror-like reflection phenomenon when submerging the SH film under water at certain tilting degrees. Lower inset compares schematically the velocity profiles under standard no-slip conditions and that over air filled pockets. The pressure was obtained from a 12 cm long channel with a height of 300 um tested at four velocities. The figure shows the pressure drop that would be expected with a smooth channel and one with a 10 μm slip on the lower surface.

In the fabricated SH film case, a mirror-like surface was observed when submerging the SH film under water at certain tilting degrees as shown inset of Fig. 6 . This reflection occurs due to the presence of fully covered air layer between the SH films and water interface44. To evaluate the impact of the SH surface, the velocity profile of the fluid flowing over the surface was obtained using micro-Particle Imaging Velocimetry (μ-PIV) techniques 45. A fluid channel was constructed that incorporated the SH surface as the lower wall and a glass coverslip as the upper wall. The average velocity of the fluid in laminar flow conditions (with a Reynolds number, Re of close to 1) was obtained across the height of the channel. The experimentally investigated conditions correspond to pressure driven Poiseuille flow conditions. The velocity profile in Fig. 6 shows the results from the surface with features of height of 30 ~ 40 μm and an RMS roughness of 9 μm, corresponding to a surface akin to that of Fig. 3(c). It was clearly seen that at the upper glass coverslip surface, the velocity tends to zero, corresponding to the expected non-slip condition. However, the velocity on the SH surface (set as the air-water interface) was non-zero (u/ umax = 0.23 from the fitted curve), with a corresponding slip length of 13μm (from Navier’s boundary condition, Vs = λ ∂u/∂y). Note that the implication of the slip length was that the liquid may be considered stagnant only at 13μm into the SH surface. The reported slip length value was one of the largest in literature43. It was then concluded that fluid flow slip does occur considerably at the SH surface, with concomitantly robust drag reducing characteristics. A more detailed interpretation of the drag reduction effects of the synthesized SH surfaces, such as their variation with Re, mean surface roughness, and lateral length scale as well as measurement of the pressure drop reduction in channel flow, is presently under investigation. In addition to the velocity profile measurements, and recognizing the uncertainty in determining the surface location for rough unpatterned surfaces, the pressure drop was measured across a larger microchannel, 12cm long by 300 μm high. The pressure drop results are shown in Fig. 6.

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The dotted lines show the pressure drop for a smooth microchannel and for one with a 10 μm slip on the lower surface. The pressure drop results shows the response is close to the 10 μm slip result.

DISCUSSION In summary, the synthesis of a mechanically robust rough surface, which seems akin to that of a bristled shark skin, exhibiting SH characteristics, through a relatively inexpensive novel roll-to-roll process has been demonstrated. The synthesis technique is simple to set up, reproducible, amenable to industrial scale production as long as we use larger size rolls, and can be adapted to widespread usage. The underlying mechanism for the formation of the rough surface has been indicated to be fingering instabilities associated with high viscosity liquids subject to roll-to-roll processes. The characterization of SH surfaces was done through static means (e.g., through contact angle measurements), as well as under dynamic/liquid flow conditions, where significant drag reduction through a pressure drop reduction was observed. The significant merit of our approach for large scale fabrication of SH surfaces is that nano-/micro-scopic patterning or chemical treatment is unnecessary for exhibiting superior SH characteristics. In addition, electrical and thermal properties may be tuned through desired electrically conductive fillers.

METHODS Raw Material Preparation and Characterization For the fabrication of the high viscosity carbon nanotube(CNT) paste, polydimethylsiloxane (PDMS, Sylgard 184 silicone elastomer base) was purchased from Dow-Corning (Midland, MI, USA), and multiwalled carbon nanotubes (MWCNTs) with an outer diameters of 10–20 nm and a lengths of 100–200 μm were purchased from Hanwa Nanotech, Inc. (Seoul, Korea). To ensure effective mixing and dispersion of the highly entangled CNTs within the polymer matrix, the pastes were premixed using a paste mixer and then a three-roll mill. MWCNTs and the PDMS base elastomer were combined in various weight ratios using the following procedure: First, the elastomer base and curing agent were mixed in a weight ratio of 10:1. This mixture was then combined with the MWCNTs in the paste mixer for 1 min.

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The resulting CNT pastes were milled for several minutes while gradually decreasing the gap between the rolls. The dynamic viscosity of the pastes was measured using a rheometer. Rheometry was conducted at room temperature in a small-amplitude oscillatory shear mode using a 20 mm parallel plate geometry with a frequency sweep from 10−1 rads−1 to 102 rads−1.

Pattern Fabrication A double-roll machine was constructed to fabricate the random SH film surfaces. The diameter and the length of the rolls were 30 mm and 300 mm, respectively, and their angular velocities were adjustable independently up to 300 rpm. A polyimide tube was placed on roller #1 as a coating target material for the CNT paste film. When the velocity of roll #1 (V1) was much higher than the velocity of roll #2 (V2), flat surface films were deposited on roll #1 when the CNT paste was added between the two rollers. At this point, inversion of the roller speeds to V1 < V2 induced capillary bridging between roll #1 and roll #2, followed by the eventual separation of the capillary bridges from roll #2 (cf. Fig. 1(b)). The optimal rolling speed ratio of roll #1 and roll #2 for generating the SH pattern depended on the viscosity of the CNT paste, the desired thickness of the sample, and so on.

Film Characterization The morphologies of the patterned films were characterized using optical microscopy and scanning electron microscopy (SEM, Quanta field emission SEM, 650 FEG). The water contact angle (WCA) and sliding angle (SA) were acquired using a contact angle measuring instrument with a drop-shape analysis system (Drop shape analyzer, KRUSS). The volume of the deionized water droplets was around 7 μl. The average contact angles were determined from a minimum of five different locations. The SA was measured at the point just before roll-off when the sample on the stage was tilted. To investigate the durability of the patterned composite films, a wear tester was constructed. Briefly, a rubber (glass) tip with a radius of 2.5 mm was applied to the film surface at a normal load of 1.5N and dragged horizontally with a sliding speed of 25 mm/s. Following the wear test, WCA measurements and SEM images were used to re-characterize the surface properties.

Velocimetry Measurement To measure the velocity profile, micro-PIV techniques were used. Fluorescent microspheres (polystyrene, diameter 1.0 μm, Bangs Laboratories, Inc)

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were distributed in distilled water and sonicated to ensure uniform particle dispersion. The flow was driven using a syringe pump through a channel (height: ~200 μm, width 9mm) at a centerline velocity of ~0.5 mm s−1, corresponding to flow in the laminar regime, Reynolds number (Re) of the order of 1. The fluid was imaged under magnification starting below the lower SH surface and past the upper non-SH surface. Upper surface can be determined by the presence of in-focus, stationary, adsorbed particles on the glass coverslip, as can the fluid/SH surface interface locations, however, due to the presence of air pockets throughout, the lower surface was set at the fluid/air/surface interface height. A set of 5 images spaced 30 ms apart were taken at each level, processed to remove background noise and outof-focus particles, and used to generate an average velocity vector for the interrogation window (field of view) per height plane. The pressure was obtained from a 12 cm long channel with a height of 300um tested at four velocities.

ACKNOWLEDGEMENTS The authors deeply appreciate the discussions and interactions with Prof BH. Kim (at University of Massachusetts Amherst).

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and spreaders. J. Fluid Mech. 7, 481–500 (1959). Owens, D. K. & Wendt, R. C. Estimation of the surface free energy of polymers. J. Appl. Polym. Sci. 13, 1741–1747 (1969). Singh, R. A., Satyanarayana, N., Kustandi, T. S. & Sinha, S. K. Tribofunctionalizing Si and SU8 materials by surface modification for application in MEMS/NEMS actuator-based devices. J. Phys. D: Appl. Phys. 44, 015301 (10pp) (2011). Bhushan, B. Biomimetics inspired surface for drag reduction and oleophobicity/philicity. Beilstein J. Nanotechnol. 2, 66–84 (2011). Rothstein, J. P. Slip on Superhydrophobic Surfaces. Annu. Rev. Fluid Mech. 42, 89–109 (2010). Lee, J. & Yong, K. Surface chemistry controlled superhydrophobic stability of W18O49 nanowire arrays submerged underwater. J. Mater. Chem. 22, 20250–20256 (2012). Meinhardt, C. D., Wereley, S. T. & Santiago, J. G. PIV measurements of a microchannel flow. Exper. Fluids. 27, 414–419 (1999).

Chapter 12

Roll-to-Roll Slot-die Coating of 400 mm Wide, Flexible, Transparent Ag Nanowire Films for Flexible Touch Screen Panels Dong-Ju Kim1,2, Hae-In Shin1, Eun-Hye Ko1, Ki-Hyun Kim3, Tae-Woong Kim3 & Han-Ki Kim1 Kyung Hee University, Department of Advanced Materials Engineering for Information and Electronics, 1 Seocheon, Yongin, Gyeonggi-do 446-701, Republic of Korea 1

Dynamic Korea Technology, R&D Center, 116-60, Sanho-daero, Gumi City, Gyeong-Buk, 39377, Republic of Korea 2

3

Samsung Display, OLED R&D Center, Yongin, Gyeonggi-do 446-711, Republic of Korea

ABSTRACT We report fabrication of large area Ag nanowire (NW) film coated using a continuous roll-to-roll (RTR) slot die coater as a viable alternative to conventional ITO electrodes for cost-effective and large-area flexible touch screen panels (TSPs). By controlling the flow rate of shear-thinning Ag NW Citation: Kim, D.-J. et al. “Roll-to-roll slot-die coating of 400 mm wide, flexible, transparent Ag nanowire films for flexible touch screen panels”. Sci. Rep.6, 34322 (2016). https://doi. org/10.1038/srep34322 Copyright © The Author(s) 2016. This work is licensed under a Creative Commons Attribution 4.0 International License. The images or other third party material in this article are included in the article’s Creative Commons license, unless indicated otherwise in the credit line; if the material is not included under the Creative Commons license, users will need to obtain permission from the license holder to reproduce the material. To view a copy of this license, visit http://creativecommons.org/licenses/by/4.0/

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ink in the slot die, we fabricated Ag NW percolating network films with different sheet resistances (30–70 Ohm/square), optical transmittance values (89–90%), and haze (0.5–1%) percentages. Outer/inner bending, twisting, and rolling tests as well as dynamic fatigue tests demonstrated that the mechanical flexibility of the slot-die coated Ag NW films was superior to that of conventional ITO films. Using diamond-shape patterned Ag NW layer electrodes (50 Ohm/square, 90% optical transmittance), we fabricated 12inch flexible film-film type and rigid glass-film-film type TSPs. Successful operation of flexible TSPs with Ag NW electrodes indicates that slot-diecoated large-area Ag NW films are promising low cost, high performance, and flexible transparent electrodes for cost-effective large-area flexible TSPs and can be substituted for ITO films, which have high sheet resistance and are brittle.

INTRODUCTION Rapid advances in flexible mobile phones, tablets, labtops, informative flat panel displays, and navigation systems for curved dash-boards in automobiles have increased demand for thin, light, transparent, highly responsive, and low-cost flexible touch screen panels (TSPs)1,2,3. Because flexible TSPs are one of the key components of high-performance flexibility displays, great effort has been focused on the development of TSPs consisting of flexible materials. To meet the advanced standards for flexible TSPs, it is imperative to develop high-quality and cost-effective transparent electrodes because the responsivity, clear visibility under various ambient light conditions, mechanical flexibility, and cost of flexible TSPs are critically dependent on the electrical, optical, and mechanical properties as well as fabrication process of transparent conductive electrodes (TCEs). The most widely used TCEs in resistive- or capacitive-type TSPs are Sn-doped In2O3 (ITO) films prepared by a vacuum-based sputtering process4,5. Although sputtered ITO films have a high conductivity and transparency in the visible region, critical problems such as the high sheet resistance of thin ITO films, the scarcity of indium resources, and the brittleness of ITO film make it impractical to use ITO films for large-area and cost-effective flexible TSPs6. In particular, the high sheet resistance (100–150 Ohm/square) of thin ITO film (20 nm) is a critical limit for realization of large-area capacitive-type TSPs above 20 inches. The need to replace high-cost and high-resistance ITO films with better performing TCE materials has yielded several TCE materials such as carbon nanotube (CNT) networks7,8, graphene sheets9,10,11, conducting

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polymer films12,13, metal nanowires14,15,16,17,18,19, as well as metal grids20,21. However, TSPs containing CNTs, graphene, and PEDOT:PSS electrodes displayed only modest performance due to the relatively high sheet resistance of CNTs and graphene and the instability of acidic PEDOT:PSS electrodes. Recently, metal grid TCEs were employed in large-area TSPs due to their low sheet resistance and high transmittance22. However, the Moiré effect, which is caused by a set of grid lines, in addition to photolithography-based complications and a high-cost patterning process are limitation of metal grid TCEs. Metal (Ag or Cu) nanowire (NW) percolating network films are being intensively investigated in academia and industry as promising TCEs for large-area flexible TSPs because of their simple and cost-effective printing process, metallic low resistivity, good flexibility, and absence of the Moiré effect23,24,25,26. Li et al., reported capacitive touch pads fabricated on paper using high-concentration Ag nanowire ink27. Lee et al., also demonstrated flexible TSPs on Mayer-rod coated PEDOT:PSS/ Ag NW hybrid electrodes28. They demonstrated the feasibility of costeffective Ag NW electrode for TSPs on non-flat surfaces27,28,29. Although the performance characteristics of Ag NW films prepared by brush painting, Mayer rod coating, spray coating, spin-coating, and a transfer process have been well studied, continuous RTR slot-die coating of Ag NW ink and application of Ag NW film on large-area flexible TSPs have not been investigated. In particular, it is important to develop a method for continuous two-step slot-die coating of Ag NWs. Furthermore, Ag NWs synthesized for use in commercialized ITO electrodes should be covered by a coating layer to protect them from sulphur in the atmosphere. In this work, we report the electrical, optical, and mechanical properties of 400 mm wide Ag NW percolating network electrodes prepared using a continuous RTR slot-die coater under atmospheric conditions. Based on our understanding of Ag NW ink flow at the lip of the slot die, we designed a slot die head appropriate for Ag NW ink and controlled the density of Ag NWs on the PET substrate. By continuous two-step RTR slot-die coating, we fabricated highly transparent and conductive Ag NW network films covered uniformly by over-coating layer to protect the Ag NWs. Capacitive-type flexible TSPs with diamond-patterned Ag NW electrodes were successfully operated, thereby demonstrating the feasibility of using the cost-effective Ag NW percolating network electrodes described here as alternative to conventional high-cost ITO electrodes.

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RESULTS Schematics of a pilot-scale RTR slot-die coating system (DKT 2015-R1SHU500) used to coat Ag NWs and an over-coating layer onto 125 μm-thick PET substrate (TORAY ADVANCED MATERIALS KOREA INC.) are provided in Fig. 1a. The RTR slot die coating system consisted of ink tanks, a slot die coating zone, a substrate heating zone, and a UV treatment zone. Using the RTR slot die coating system, the Ag NW layer was uniformly coated on the PET substrate through the slot die head, and then the films passed through the heating zone (120 °C) by means of unwinding and rewinding the roller at a constant speed of 2 m/min. Figure 1b exhibits slot die coating of Ag NWs on the moving PET substrate at room temperature. After coating of the Ag NWs, the over coating layer was also uniformly coated on the Ag NW layers. Ag NWs covered by the over-coating layer then entered the heating zone (80 °C) and UV treatment zone filled with nitrogen ambient by means of unwinding and rewinding the roller at a constant speed of 2 m/min. The purpose of the over-coating layer was to impart mechanical strength to the films and protect the Ag NWs layer from direct environmental exposure. As discussed by Yacaman, atmospheric corrosion of Ag NWs resulted in the formation of silver sulphide nanocrystals on the surfaces of Ag NWs30. Therefore, an effective over-coat layer is to protect against atmospheric corrosion to enable commercialization of slot-die-coated Ag NW network electrodes. For simplicity, we refer to the Ag NWs covered by the overcoating layer as OC-Ag NW films hereafter. Figure 1c shows the in-situ measured sheet resistance of OC-Ag NW films fabricated using a RTR slotdie coater as a function of pump frequency used to feed the Ag NW ink to control the density of the Ag NW network.

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Figure 1: (a) Schematic illustration of the continuous RTR slot-die coating system used to coat Ag NWs and over coating layer on the PET substrate. The RTR slot-die system consisted of a slot-die coating zone, heating zone, and UV treatment zone. (b) Schematics of slot-die coating of Ag NWs and the over-coating layer. (c) Picture of Ag NW percolating network electrodes with different sheet resistances (30 to 70 Ohm/square).

To fabricate large area TSPs and confirm the operation uniformity of flexible TSPs, uniform coating of Ag NW ink is very important. Therefore, proper design of the slot die to unsure uniform distribution of Ag NW ink is critical for the fabrication of large area Ag NW network films with uniform

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sheet resistance. Figure 2a shows a picture and the schematic structure of the slot die head in the RTR slot-die coating system. In the slot die coating process, the Ag NW ink is pumped from the Ag NW ink tank to a die and distributed across the width of a narrow slot. The Ag NW ink filled the gap between the die lips and moving PET substrate and formed Ag NW network films. The thickness of the Ag NW film was controlled by the Ag NW ink flow rate to the slot die and the speed of the moving PET substrate. Right schematic in Fig. 2a illustrates the flow of Ag NW ink at the exit of the slot die. To optimize the design of the slot-die head for Ag NW ink, we measured the viscosity of Ag NW ink as a function of shear rate. The viscosity and shear stress of Ag NW ink as a function of shear rate were analyzed using Rheoplus Software as shown in Fig. 2b. The viscosity of Ag NW ink decreased with increasing shear rate (shear thinning)31,32. Therefore, the ink acted as a non-Newtonian fluid. The uniformity of the slot-die coated Ag NW ink was critically affected by the PET substrate moving speed, the viscosity of the Ag NW ink, and the geometry of the upstream and downstream lips. Based on use of Ag NW ink with non-Newtonian behaviour, we designed a slot die head for our pilot-scale RTR slot die coater. Under optimized slot-die coating and substrate heating conditions, we achieved uniformly coated Ag NW network films with different sheet resistances as shown in Fig. 2c. In general, the sheet resistance of Ag NWs network is closely related to the density of Ag NWs. Therefore, an increase in the pumping frequency of Ag NW ink into the slot die head led to a decrease in sheet resistance from 70 to 30 Ohm/square, as shown in Table 1. Figure 2d shows a picture of 400 mm wide OC-Ag NW network films obtained after continuous two-step slot die coating. Sheet resistance was measured in-situ in the direction of width. Due to the uniform distribution of Ag NWs in the lips and the constant PET velocity and tension, the RTR slot-die coated Ag NW network films on the 400 mm-wide PET substrate showed good uniformity regardless of position. The inset picture in Fig. 2d shows the in-situ sheet resistance of a RTR slotdie-coated Ag NW network electrode with a sheet resistance of 50 Ohm/ square.

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Figure 2: (a) Picture of the slot die head ejecting Ag NW inks on the PET substrate and schematic structure of a slot die head used for Ag NW coating. (b) Viscosity and shear stress of Ag NWs ink versus shear rate. (c) Sheet resistance of the slot-die coated Ag NW network electrode as a function of increasing pump frequency for feeding Ag NW ink. (d) Picture of OC-Ag NW films and sheet resistance of Ag NWs at different positions indicating good uniformity of the Ag NW network films.

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Table 1: Electrical and optical properties of slot-die coated OC-Ag NW films as function of pumping frequency to control the density of Ag NW network

Surface FESEM images of slot-die coated Ag NW network films as a function of pumping frequency are shown in Fig. 3a. It is clear from these images that the coverage of Ag NWs on the PET substrate was mainly affected by the pumping frequency used to feed the Ag NW ink into the slot die. As expected from the sheet resistance of Ag NW films in Fig. 2c, an increase in pumping frequency led to an increase in percolating Ag NW density and a decrease in sheet resistance of the Ag NW network films. Figures 3b show the optical transmittance the OC-Ag NW network films with different sheet resistances ranging from 30 to 70 Ohm/square. All RTR slot-died coated Ag NW network films had a high optical transmittance and low reflectance in the visible wavelength region, as well as a broad surface plasmon resonance band in the UV region33. Compared to OC-Ag NW network films with sheet resistances of 50 and 70 Ohm/square, respectively, the Ag NW network film with a sheet resistance of 30 Ohm/square showed slightly lower optical transmittance due to scattering by the high-density Ag NWs. OC-Ag NW films with a sheet resistance of 70 Ohm/square had the highest transmittance of 90% and the lowest reflection of 10% at a wavelength of 550 nm. Although the increase in Ag NW density led to a decrease in optical transmittance, all RTR slot-die coated Ag NW network films exhibited optical transmittance in the visible wavelength region that was high enough for use in flexible TSPs. In addition, the optical transmittance of the Ag NW networks was higher than that of semi-transparent Ag film, because the high transmittance of Ag NWs is achieved by passing of light through the Ag NW uncovered region. As shown in Fig. 3c, the low reflectance of the Ag NW network regardless of the sheet resistance indicates that light reflection by Ag NWs is negligible due to their very narrow width (~22 nm). Figure 3d shows haze and optical transmittance values of the Ag NW network films with different sheet resistances. Haze of TCEs is defined as the percentage of transmitted light passing through the film that deviates more than 2.5° from the incident beam by forward scattering1:

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Figure 3: (a) Surface FESEM images of RTR slot-die coated Ag NW networks as a function of pumping frequency. (b) Optical transmittance and (c) reflectance of RTR slot-die-coated Ag NW network films with decreasing Ag NW densities. (d) Haze and transmittance at 550 nm wavelength region of RTR slotdie-coated Ag NW network films. (e) Optical transmittance and reflectance of diamond-patterned Ag NW film.

(1) Here, Id is the light flux transmitted directly, and (Is)f is the flux that undergoes forward scattering, i.e. scattering intensity between 0° and 90°. As discussed by Mini et al., TCEs should have a low haze value for successful integration of flexible TSPs into flexible displays, because the blurriness of TSPs is mainly related to the haze value of the TCE film17. The recommended haze

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value of TCEs used in high performance TSPs is less than 1%. The diamondshape patterned Ag NW electrodes showed high optical transmittance and low reflectance in addition to acceptable haze values. As shown in Table 1 and Fig. 3d, the haze of the OC-Ag NWs layer is lower than that of pure Ag network due to the existence of the over coating layer, which uniformly covered the slot-die coated Ag NWs network and improved the flatness of the OC-Ag NWs network. Therefore, reduced light scattering on the OCAg NWs decreased the haze value of the OC-Ag NWs network films17,22. In addition, due to the high optical transmittance of patterned OC-Ag NW network films as shown in Fig. 3e, flexible TSPs consisting of merged Ag NW network films showed a high an optical transmittance of 90% and a low reflectance of 10%. Table 1 summarized the electrical and optical properties of slot-die coated Ag NW network film as a function of pump frequency. Cross-sectional TEM images of RTR slot-die-coated Ag NW network film with a sheet resistance of 50 Ohm/square on PET substrate in addition to an XRD plot insert are shown in Fig. 4a. The slot-die coated Ag NWs were randomly distributed on the PET substrate with smooth surface morphology. As shown in the XRD plot in inset of Fig. 4a, the Ag NWs were terminated by the (111) plane and the (100) side surface plane. Both the (111) end plane and (100) side surface plane of the Ag NWs were physically connected and provided conduction paths for electrons. In particular, the slot-die coated Ag NWs were parallel to the PET substrate. The enlarged TEM image of a cross-section of an Ag NW in Fig. 4b shows the circle-shaped end of an Ag NW and a pentagonal rod with five boundaries, as indicated by arrows. As discussed by Chen et al., anomalous contrast fringes in boundaries indicate the presence of stacking faults and twinning boundaries34. Cross-sectional images of single Ag NW with diameters of around 18.7 around 22.6 nm, respectively, are shown in Fig. 4c. The enlarged images shows that the single Ag nanorod had a round-shaped end. Figure 5 shows TEM images obtained from Ag NW-Ag NW junction regions. As shown in Fig. 5a,b, the junctions between Ag NWs were cross-bar and laterally parallel junctions. The crosssectional HRTEM image revealed that the (100) side surface plane of the Ag NW contacted the other (100) side surface plane of the Ag NW. Considering percolating Ag NW networks and the optical transparency of the uncovered Ag NW region, a cross-bar junction is more advantageous for decreasing the sheet resistance of Ag NW film, because the cross-bar junctions of Ag NWs provide effective conduction paths for randomly distributed Ag NWs. The enlarged TEM image in Fig. 5c shows that the (100) surface of the Ag NW side plane was well connected to other surface (100) plane of the Ag NW.

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Figure 4: (a) Cross-sectional TEM image of RTR a slot-die coated Ag NW network on PET substrate. (b) High resolution TEM image of the end of a single Ag nanowire showing five distinct boundaries. (c) Enlarged TEM image of the side plane of Ag nanowires with diameters of around 18.7 and 22.6 nm, respectively.

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Figure 5: Cross-sectional TEM image of (a) a cross-bar junction and (b) laterally parallel junction between Ag NWs. (c) High resolution TEM image of the (100)-(100) side plane junction between Ag NWs.

To demonstrate the feasibility of Ag NW network films as flexible TCEs for flexible TSPs, we investigated their mechanical flexibility based on specially designed bending test systems. The left pictures in Fig. 6ashows the outer/inner bending systems that we used to measure changes in resistance of various outer/inner bending radii. As shown on the right in Fig. 6a, the change in resistance of an Ag NW network film can be expressed as (RR0)/R0, where R0 is the initial measured resistance and R is the resistance measured under substrate outer/inner bending35,36. The outer bending test

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results revealed that the slot-die coated Ag NW network film kept a constant resistance until the bending radius reached 2.5 mm. However, a further decrease in the outer bending radius rapidly increased the resistance changes due to tearing of the OC-Ag NW layer, as shown in Fig. 6b. Compared to conventional ITO film with a critical outer bending radius of 8–10 mm, the slot-die coated Ag NW network film showed superior flexibility due to the high strain failure of Ag. In the inner bending tests, the Ag network film also showed a constant resistance change until the sample was bent to an inner bending radius of 2 mm (the bending limit). Even though the OC-Ag NW network films delaminated from the PET substrate or many cracks formed in the OC-Ag NW network films under severe inner bending, the change in resistance was smaller than that observed during the outer bending test. Under inner bending, the flexible Ag network film showed small resistance changes because of the overlapping of cracked or delaminated layers. However, when outer bending was applied, the Ag network films were under tensile stress, as shown in the inset of Fig. 6a. Due to this tensile stress, cracks isolated the OC-Ag network and increased the resistance change when it was severely bent below the bending radius of 7 mm, as shown in the surface FESEM image of the OC-Ag network after a 3 mm outer bending test. Cracks parallel to the outer bending direction separated the OC-Ag network films. Figure 6c shows the dynamic outer and inner bending test results of OC-Ag network films as a function of increasing bending cycles for a fixed outer and inner bending radius of 5 mm, which is the requested bending radius in flexible TSPs. Both dynamic outer and inner bending fatigue tests showed no change in resistance (ΔR) after 10,000 bending cycles, demonstrating the good mechanical flexibility of the slot-die coated Ag network films. Figure 6d shows resistance changes of the OC-Ag network film during a twisting test at a fixed twisting angle of 10°. The OC-Ag network film showed constant resistance as the number of twisting cycles increased, indicating that the slot-die coated OC-Ag network films had good flexibility. Figure 6eshows the resistance changes of the OC-Ag network film during the rolling test. The OC-Ag network film was clipped on a rolling bar with a radius of 70 mm as shown in the inset picture, and then rolled repeatedly. During a 10,000 cycle rolling test, the OC-Ag network film showed constant resistance changes. Based on the results of the outer/inner, twisting, and rolling tests, we concluded that the RTR slot-die coated OC-Ag network film had sufficient flexibility to enable realization of flexible TSPs

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Figure 6: (a) Picture of the outer and inner bending test system and resistance changes during outer and inner bending as a function of bending radius. (b) Surface FESEM image of OC-Ag NW network film after a 3 mm radius outer bending test. Arrows indicate the regions separated by cracks. (d) Dynamic outer and inner bending fatigue tests of OC-Ag NW network films at a fixed bending radius of 5 mm. (d) Twisting test of the OC-Ag network film at a twist angle of 10° with inset picture showing twisting steps. (e) Rolling test of OCAg network films with increasing rolling cycles; the inset picture shows the rolling step.

Figure 7a shows the schematic patterning process of RTR slot-die-coated OC-Ag network film based on a wet etching system. By coating a liquid photo resist (LPR) layer and then exposing the positive-masked LPR/OC-

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Ag NW network/PET films to UV, we successfully patterned OC-Ag NW network films with a typical diamond shape. Optical microscope images of the diamond shape-patterned bottom of Ag NW network films are shown in Fig. 7b. Detailed wet-etching process of the Ag NW film for diamondpatterning was explained.

Figure 7: (a) Schematic of pattering process of RTR slot-died coated Ag network films. (b) Optical microscope images of the diamond-shape patterned Ag network films.

A schematic structure of flexible TSPs with diamond-shaped top and bottom Ag NW network films is shown in Fig. 8a. By merging the diamondpatterned bottom and top Ag NW network films, we fabricated flexible TSPs with the structure of PET/top Ag NW film/OC-OC/bottom Ag NW film/ PET. In general, capacitive-type TSPs are built using two TCEs in parallel, where the patterns of the TCEs form a capacitor holding charge. Touching a finger to the TSPs changes the electric field between the capacitor plates, because human body capacitance absorbs the fringing electric field, as shown in Fig. 8a. Pictures of 12-inch flexible TSPs fabricated on patterned Ag NW electrodes with a low sheet resistance of 50 Ohm/ square and high optical transmittance of 90.0% are shown in Fig. 8b. Due to the high optical transmittance of patterned OC-Ag NW network films, flexible TSPs consisting of merged Ag NW network films showed high an optical transmittance as well as good flexibility. By connecting the TSPs to software (MS paint program), we were able to operate the flexible TSP based on the diamond-patterned Ag NW network films. Figure 8c shows the writing function of the flexible TSPs based on diamond-patterned Ag NW network films. Generally, GFF-type TSPs operates by exact sensing of X-Y coordinates and the characteristics of linearity. TSPs with a diamond-shaped

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Ag NW network films were also operated with protective cover glass. TSPs with diamond-patterned Ag NW network films were successfully used to perform writing functions. This demonstrated that the diamond-patterned Ag NW network films, which had low sheet resistance and high optical transmittance as well as good mechanical flexibility, are a promising transparent, flexible, and cost-effective electrodes that can substitute for conventional ITO electrodes in large area flexible TSPs.

Figure 8: (a) Schematics of flexible TSP operation with RTR slot-die coated Ag NW network films. (b) Picture of 12-inch curved TSPs with RTR slot-die coated Ag NW network films. (c) Operation of TSPs with an Ag NW network electrode.

CONCLUSION In summary, we developed a simple RTR slot die coated Ag NW network films for large area flexible TSPs and demonstrated the feasibility of Ag

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NW electrodes as promising replacements for conventional ITO electrodes. Based on an understanding of Ag NW ink flow and proper slot-die design, we fabricated 400 mm-wide Ag NW network films with a sheet resistance uniformity within 5% and controlled the density of Ag NWs by controlling the injection flow of Ag NW ink. The tunable electrical and optical properties of the Ag NW network films were affected by the pumping frequency used to inject Ag NW ink into the slot die. Large-area Ag NW network films with low sheet resistance and high optical transmittance produced by this process are a viable alternative to high-cost ITO films prepared by high-cost vacuum-based sputtering processes. By using diamond-shape patterned Ag NW network films, we obtained highly transparent Ag NW electrodes with a low sheet resistance of 50 Ohm/square, which is acceptable for fabrication of large-area TSPs. Outer/inner, twisting, and rolling test results demonstrated the superior mechanical flexibility of the Ag NW network film to that of conventional ITO films. We further demonstrated the writing operation of flexible TSPs using diamond-patterned Ag NW network electrodes. Together, our results indicate that RTR slot-die-coated Ag NW networks are promising substitutes for conventional ITO electrodes in large-area and low cost flexible TSPs.

METHODS Roll-to-roll slot-die Coating of Ag NWs and the Over Layer 400 mm-wide Ag NW network films were prepared on a PET substrate (TORAY ADVANCED MATERIALS KOREA INC.) at room temperature using a specially designed pilot-scale RTR slot-die coating system. The RTR slot die coating system(DKT 2015-R1-SHU500) consisted of ink tanks, a slot die coating zone, a substrate heating zone, and a UV treatment zone. Using the RTR slot die coating system, an Ag NW layer was uniformly coated on a PET substrate through the slot die head and then films were moved to the heating zone (120 °C) by means of unwinding and rewinding at a roller speed of 2 m/min. After coating the Ag NWs, an over-coating layer was uniformly coated on the Ag NW layers to protect the Ag NWs and prevent their degradation. The Ag NWs covered by an over-coating layer were then moved to the heating zone (80 °C) and the UV treatment zone filled with nitrogen ambient.

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Characterization of RTR slot-die Coated Ag NW Network Films The sheet resistance and resistivity of the Ag NW network films were measured by Hall measurements (HL5500PC, Strength 0.32 T, Accent Optical Technology) at room temperature. The sheet resistance also were measured by tip radius 700 μm of 4-point probe (Model : RSP-1000, DASOLENG in Korea). The optical transmittance of the Ag NW network films was measured by UV/visible spectrometry (Lambda 35) carried out over the wavelength range from 220 to 800 nm. The structural properties of Ag NW network films were analyzed by means of X–ray diffraction (D/MAX-2500). The microstructures and contact region of the Ag NW network electrodes were examined by high-resolution electron microscopy (HRTEM:S-4800Hitachi). Cross-sectional HREM specimen was prepared by means of focus ion beam (FIB) milling. The mechanical properties of OC-Ag NW network films were evaluated using a specially designed inner/outer bending, twisting, and rolling system. In addition, dynamic fatigue bending tests were carried out using a lab-designed cyclic bending test machine, operated at the frequency of 0.5 Hz for 10,000 cycles. The resistances of the Ag NW network films were measured throughout cyclic bending. Furthermore, twisting and rolling tests were carried out using labdesigned twisting and rolling test system at a constant twisting angle of 10° and rolling radius of 70 mm.

Patterning of Ag NW Network Films and Fabrication of Flexible TSPs Prior to the diamond patterning of OC-Ag NWs films, the OC-Ag NWs films were annealed in oven box at 130 °C for 20 min to prevent film shrinkage. Then, a liquid photo resist (LPR : AZ HKT-601) layer was coated onto the OC-Ag NWs films by a spin-coater. Then, the LPR-coated OCAg NW films were exposed to UV light at 60 mJ using a positive diamond mask. The UV-exposed OC-Ag NW films were patterned using developing solution (EN-DT238E: tetramethylammonium hydroxide 3%, surfactant 2%, deionized water 95%). The diamond-patterned OC-Ag NW films were subsequently etched using an etching solution (EO- NS100: nitric acid & deionized water). Finally, the stripped OC-Ag NWs films were cleaned by a spray-type rinse system using deionized water. To fabricate flexible TSPs, Ag paste was directly printed on the diamond-patterned OC-Ag NW film by silk screen printing system. The resulting OC-Ag NW film/OCA/OCAg NW film was connected to a flexible printed circuit board by bonding

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both the metal pattern and the FPCB to an anisotropic conductive film. Both top OC-Ag NW/PET and bottom OC-Ag NW/PET films were bonded to a flexible printed circuit board (FBCB) using an anisotropic conductive film, and then cover glass was attached to the top OC-Ag NW/PET film using OCA film to protect the diamond-patterned top OC-Ag NWs/PET film. Finally, the FPCB was connected to an IC controller to operate the TSPs.

ACKNOWLEDGEMENTS The authors acknowledge the financial support of the National Research Foundation of Korea (NRF) grant funded by the Ministry of Education, Science and Technology (NRF 2015R1A2A2A01002415). This study also received partial support from Samsung Displays. LTD.

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Ye, S. et al. Metal Nanowire Networks: The Next Generation of Transparent Conductor. Adv. Mater. 26, 6670 (2014). Ghaffarzadeh, K. & Das, R. Transparent conductive films (TCF) 2015– 2025: Forecasts, Markets, Technologies (IDTechEx), pp. 2014-2042. Jeong, C. W. et al. Prospects for nanowire-doped polycrystalline graphene films for ultratransparent, highly conductive electrodes. Nano Lett. 11, 5020 (2011). Jung, H. K., Kim, J. T., Sahama, T. & Yang, C. H. Future Information Communication Technology and Applications. Springer-Volume 1, 445 (2013). Cairns, D. R. et al. Strain-dependent electrical resistance of tin-doped indium oxide on polymer substrates. Appl. Phys. Lett. 76, 1425 (2000). Kumar A. & Zhou, C. The race to replace tin-doped indium oxide: which material will win? ACS Nano 4, 11 (2010). Hecht, D. S., Hu, L. & Irvin, G. Emerging Transparent Electrodes Based on Thin Films of Carbon Nanotubes, Graphene, and Metallic Nanostructures. Adv. Mater. 23, 1482 (2011). Feng, C. et al. Flexible, stretchable, transparent conducting films made from superaligned carbon nanotubes. Adv. Funct. Mater. 20, 885 (2010). Kim, K. S. et al. Large-scale pattern growth of graphene films for stretchable transparent electrodes. Nature 457, 706 (2009). Lee, J. S., Novoselov, K. S. & Shin, H. S. Interaction between metal and graphene: dependence on the layer number of graphene. ACS Nano 5, 608 (2011). Bae, S. K. et al. Roll-to-roll production of 30-inch graphene films for transparent electrodes. Nature Nanotechnol. 5, 574 (2010). Kirchmeyer, S. & Reuter, K. Scientific importance, properties and growing applications of poly(3,4-ethlenedioxythiophene). J. Mater. Chem. 15, 2077 (2005). Argun, A. A., Cirpan, A. & Reynolds, J. R. The first truly all polymer electrochromic devices. Adv. Mater. 15, 1338 (2003). De, S. et al. Silver networks as flexible, transparent, conducting films: extremely high dc to optical conductivity ratios. ACS Nano3, 1767 (2009).

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15. Xu, F. & Zhu, Y. Highly conductive and stretchable silver nanowire conductors. Adv. Mater. 24, 5117 (2012). 16. Hu, L., Kim, H. S., Lee, J. Y., Peumans, P. & Cui, Y. Scalable Coating and Properties of Transparent, Flexible, Silver Nanowire Electrode. ACS Nano 4, 2955 (2010). 17. Menamparambath, M. M. et al. Silver nanowires decorated with silver nanoparticles for low-haze flexible transparent conductive films. Sci. Rep. 5, 16371 (2015). 18. Kang, M.-G., Park, H.-J., Ahn, S. H. & Guo, L. Jay. Transparent Cu nanowire mesh electrode on flexible substrates fabricated by transfer printing and its application in organic solar cells. Sol. Energy Mater. Sol. Cells 94, 1179 (2010). 19. Lee, J. Y., Connor, S. T., Cui, Y. & Peumans, P. Solution-processed metal nanowire mesh transparent electrodes. Nano Lett. 8, 689 (2008). 20. Kang, M. G., Kim, M. S., Kim, J. S. & Guo, L. J. Organic solar cells using nanoimprinted transparent metal electrodes. Adv. Mater.20, 4408 (2008). 21. Tvingstedt, K. & Inganas, O. Electrode grids for ITO-free organic photovoltaic devices. Adv. Mater. 19, 2893 (2007). 22. Mohl, M., Dombovari,A., Vajtai, R., Ajayan,P. M. & Kordas,K. Selfassembled large scale metal alloy grid patterns as flexible transparent conductive layers. Sci.Rep. 5, 13710 (2015). 23. Liu, C. H. & Yu, X. Silver nanowire-based transparent, flexible, and conductive thin film. Nanoscale Res.Lett. 6, 75 (2011). 24. Rathmell, A. R. & Wiley, B. J. The synthesis and coating of long, thin copper nanowires to make flexible, transparent conducting films on plastic substrate. Adv. Mater. 23, 4798 (2011). 25. Zhang, D. et al. Synthesis of ultralong copper nanowires for highperformance transparent electrode. J. Am. Chem. Soc. 134, 14283 (2012). 26. Kim, D. J. et al. Indium-free, highly transparent, flexible Cu2O/Cu/ Cu2O mesh electrodes for flexible touch screen panels. Sci. Rep. 5, 16838 (2015) 27. Li, R.-Z., Hu, A., Zhang, T. & Oakes, K. D. Direct writing on paper of foldable capacitive touch pads with silver nanowire inks. ACS Appl. Mater. & Inter. 6, 21721(2014). 28. Lee. J. H. et al. Room-temperature nanosoldering of a very long metal

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nanowire network by conducting polymer assisted joining for a flexible touch-panel application, Adv. Fun. Mater. 23, 4171 (2013). Lee, J. H. et al. Very long Ag nanowire synthesis and its application in a highly transparent, conductive and flexible metal electrode touch panel. Nanoscale 4, 6408 (2012). Elechiguerra, J. L. et al. Corrosion at the Nanoscale: The case of silver nanowires and nanoparticles. Chem. Mater. 17, 6042 (2005). Romero, O. J., Suszynski, W. J., Scriven, L. E. & Carvalho, M. S.Lowflow limit in slot coating of dilute solutions of high molecular weight polymer. J. Non-Newtonian Fluid Mech. 118, 137 (2004). Jiang, B., Keffer, D. J., Edwards, B. J. & Allred, J. N. Modeling shear thickening in dilute polymer solutions: temperature, concentration, and molecular weight dependencies. J. Appl. Poly. Sci. 90, 2997 (2003). Kottmann, J. P., Martin, O. J. F., Smith, D. R. & Schultz, S. Plasmon resonances of silver nanowire with a nonregular cross section. Physical Review B. 64, 235402 (2001). Chen, H. et al. Transmission-Electron-Microscopy Study on Fivefold Twinned Sliver Nanorods. J. Phys. Chem. B. 180, 12038 (2004). Seo, K. W., Lee, J. H., Kim, H. J., Na, S. I. & Kim, H. K. Highly transparent and flexible InTiO/Ag nanowire/InTiO films for flexible organic solar cells. Appl. Phys. Lett. 105, 031911 (2014). Lee, J. H. et al. Brush painting of transparent PEDOT/Ag nanowire/ PEDOT multilayer electrodes for flexible organic solar cells. Sol. Energy Mater. Sol. Cells 114, 15 (2013).

SECTION V: SURFACE TENSION DRIVEN DEFECTS AND STATIC ELECTRICITY

Chapter 13

Surface Tension-Driven Self-Alignment

Massimo Mastrangelia, Quan Zhoub, Veikko Sariolac, and Pierre Lambertd Physical Intelligence Department, Max Planck Institute for Intelligent Systems, Max Planck ETH Center for Learning Systems, Heisenbergstr. 3, 70569 Stuttgart, Germany a

b Department of Electrical Engineering and Automation, School of Electrical Engineering, Aalto University, Otaniementie 17, 02150 Espoo, Finland

Department of Automation Science and Engineering, Tampere University of Technology, Korkeakoulunkatu 3, 33720 Tampere, Finland c

Department of Bio, Electro And Mechanical Systems, E´cole Polytechnique de Bruxelles, Universite´ Libre de Bruxelles, CP 165/56. Avenue F.D. Roosevelt 50, 1050 Brussels, Belgium d

ABSTRACT Surface tension-driven self-alignment is a passive and highly-accurate positioning mechanism that can significantly simplify and enhance

Citation: M. Mastrangeli, Q. Zhou, V. Sariola and P. Lambert, “Surface tension-driven selfalignment” Soft Matter, 2017, 13, 304 https://dx.doi.org/10.1039/c6sm02078j Copyright © 2017 by the authors. This article is licensed under a Creative Commons Attribution 3.0 Unported Licence. Material from this article can be used in other publications provided that the correct acknowledgement is given with the reproduced material.

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the construction of advanced microsystems. After years of research, demonstrations and developments, the surface engineering and manufacturing technology enabling capillary self-alignment has achieved a degree of maturity conducive to a successful transfer to industrial practice. In view of this transition, a broad and accessible review of the physics, material science and applications of capillary self-alignment is presented. Statics and dynamics of the self-aligning action of deformed liquid bridges are explained through simple models and experiments, and all fundamental aspects of surface patterning and conditioning, of choice, deposition and confinement of liquids, and of component feeding and interconnection to substrates are illustrated through relevant applications in micro- and nanotechnology. A final outline addresses remaining challenges and additional extensions envisioned to further spread the use and fully exploit the potential of the technique.

INTRODUCTION Understanding the role that surface tension plays in the multiform interactions between liquid phases and solid surfaces drives a significant part of current research in fluidics,1mechanics,2 soft matter3 and interfacial engineering.4 Practical implications of new insights gained from basic research can rapidly lead to applications in wetting and surface conditioning,5 precision engineering,6 colloidal assembly,7 microfluidics8 and micro/ nanofabrication.9 In particular, the accurate and repeatable manipulation of micro- and nanoscopic objects represents a technological field of farreaching industrial, biomedical and commercial relevance that is vastly benefiting from such research.10 In this context, for the past two decades fluid-mediated manipulation has demonstrated to be capable of simplifying and enhancing the construction of heterogeneous micro- and nanosystems. Precise positioning is of foremost importance when assembling together sets of modules or devices that are functionally correlated. Recent advances in miniaturization, especially in the pitch and diameter of interconnections, and in the integration of diverse complementary functions often make the realization of novel microelectronic, optoelectronic and nanoscale systems possible only through extremely accurate registration among the modules. Robotic assembly10 and stochastic self-assembly9represent the opposite extremes of a range of approaches to precision assembly.11Robotic pick-andplace is the current standard in precision industrial assembly. It achieves impressive performance (on the order of tens of thousands of assembled

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components per hour in the case of flip-chip assembly machines12) through the supervised dexterity of robotic effectors. The throughput of robotic pickand-place is however limited by its seriality, while its accuracy degrades when handling components of sub-millimetric size.13 Self-assembly is an alternative, bio-inspired and intrinsically parallel technique for highthroughput manufacturing,12 particularly for handling very small, ultrathin and flexible components. Surface tension-driven (alias capillary) selfalignment (Fig. 1) lies somewhat in the middle of the assembly spectrum, as it merges deterministic aspects of component handling and feeding with passive, unsupervised and very accurate component registration. When compared to stochastic fluidic self-assembly, the key advantage of capillary self-alignment is that its steps can more easily be individuated, parameterized and streamlined into industrial process flows.

Figure 1: Capillary self-alignment. Top row: Sketch of the process. (a) A component is brought in the vicinity of a droplet confined on a receptor site at offsets x0, y0 (not shown) and z0 with respect to its final position; (b) description of the process based on the minimization of the surface energy (see Section 2.3); (c) dual description based on a capillary force vector acting on the component (Section 2.4); (d) end of the alignment process, with possible residual misalignment. Bottom row: Sequential snapshots of the capillary selfalignment of a square transparent plastic component onto a shape-matching receptor coated with a thin water film. The autonomous roto-translational motion of the component starts upon its release from the tweezer only after the triple contact line reaches at least one of the solid edges of the bottom surface

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of the component. Final placement of the component on the receptor and alignment of the edges is achieved after a few oscillation cycles (not shown) due to underdamped dynamics. (Adapted with permission from ref. 20 © 2014, American Chemical Society.).

The development of capillary self-alignment progressed across three subsequent phases. In the first phase, the surface tension of molten solderbased interconnections was used to self-align and electromechanically fix surface-mount components to electronic substrates – a process originally introduced in the ‘60s14 which came to be generally associated with flipchip assembly. Through continuous evolution, including its adaptation to three-dimensional capillary deployment of microelectromechanical systems (MEMS) 15 and capillary self-folding of micro- and nanoscopic voxels,16presently molten solder-based self-alignment is amply used in advanced microelectronic integration processes involving solder microbumps, through-silicon vertical interconnect accesses (VIAs) and three-dimensional chip stacking to meet ultimate performance and input/ output requirements. The second phase, started in the late ‘90s, saw the development of a multitude of fluidic self-assembly approaches. Among them, capillarity-driven component-to-substrate stochastic self-assembly emerged as the most convenient for industrial adoption. In the third phase and during the last decade, the fluidic self-assembly process was adapted to work in an air environment. This evolution is akin to a return to the origin and enabled the combination with robotic component feeding into what was called hybrid microhandling.17 Such latest embodiment indeed consists of an extension of the IBM›s original controlled-collapse chip connection (C4) integration process,14 whereby an additional liquid bridge wraps and complements the function of the solder bumps.18 The outlined stages of development tackled advances in technology and materials science, among which in particular surface conditioning and patterning, wetting control, new interconnection materials, ultrathin and flexible components. While most of past and current research concerned inorganic components, capillary self-alignment may hold potential also for organic assemblies. The understanding of the impact of process parameters – such as surface tension, viscosity and contact angle of liquids, geometry of components and receptors, relative misalignment – also gradually improved thanks to the development of analytical and numerical models. While the theoretical picture of capillary self-alignment, especially regarding its dynamics, is not fully clear yet, the available frameworks have contributed to increase assembly throughput and yield through better process design.

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This review is tailored to guide the readers through the practice of capillary self-alignment. The underlying physics and the material requirements of the process are described together with the state-of-the-art in its performance and achievements. The paper is intended as a tutorial to help transpose prior experimental demonstrations into industrial practice. Recent works19 and patents support the claim that capillary self-alignment›s current state of development is compatible with this forthcoming step. The paper is structured as follows. Definitions and fundamental concepts concerning wetting and triple contact lines are first summarized to introduce an overview and the governing physics of surface tension-driven self-alignment (Section 2). The design of constrained liquid bridges is then addressed, highlighting the importance of surface properties in the methods of liquid confinement (Section 3) and the role of shape matching between component and receptor (Section 4). Section 5 provides an overview of the dynamics of capillary self-alignment, as reflected in models and experimental data. Process integration is discussed in the following Section 6, concerning the choice of liquids, techniques for controlled liquid deposition and feeding of components to the receptors, and available options to electrically besides mechanically bind components to substrates upon completion of capillary self-alignment. Relevant applications of surface tension-driven selfalignment reported in recent literature are also exemplified, and a survey of related patents is provided in Section 7. Open challenges and future perspectives are outlined in the concluding Section 8.

FUNDAMENTALS OF SURFACE-TENSION-DRIVEN SELF-ALIGNMENT Definitions and Terminology Surface-tension-driven self-alignment (also known as capillary selfalignment) is a fluid-based technology used to accurately align components to patterns or features of a substrate. When a confined liquid drop is interposed between a suitably patterned region of a substrate and a freely moving component, the droplet forms a liquid meniscus whose constrained capillary action tends to align the component to the pattern. The attribute “self” refers to the apparent autonomy of the transient component motion. A sketch of the process and sequential snapshots from a capillary selfalignment experiment are shown in Fig. 1.

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The object to be assembled – typically a chip, die, device or functional module in electronics applications – is here called component (C). The well-defined area ALR, on a substrate or on another component, to which the component is required to self-align represents the binding, assembly or target site, hereby referred to as receptor (R). The finite liquid phase L, bounded by the gas phase G and interposed between the component and the receptor to form a liquid bridge or meniscus,4 additionally prevents dry friction between the sliding surfaces of component and receptor and hence is sometimes called lubricant. The height h of the liquid bridge defines the gap. Most of the emphasis in capillary self-alignment is reserved to in-plane motion, i.e. the biaxial (x, y) and rotational (θz) motion of the component relative to the receptor. This represents a simplified account, as the alignment process inevitably involves perturbations and motion of the component along all six degrees of freedom of the fluid joint,2,4 whereby vertical motion (z) and out-of-plane rotations (θx and θy) are also included. The emphasis on lateral registration derives from technological applications, such as precision assembly tasks and micro/nanoelectronic integration, and has in turn largely biased the study and modelling of the process. In terms of component placement, the bias or offset refers to the initial distance (i.e.prior to the inception of the self-aligning motion) of the component to its final expected position – typically the one where the edges of the component are perfectly aligned to those of the receptor. The offset vector identifies the component position at the time the component is released, in case of robotic feeding, or when the component is first wet by the droplet, in case of stochastic feeding (Fig. 1a). Misalignment is conversely a measure of registration error, and refers to the eventual residual distance between actual and expected position of the component after completion of the self-alignment process (Fig. 1d, top row). Misalignment may arise from wetting defects (Section 3.2) and shape mismatch (Section 4.2). Null misalignment identifies correct final component positioning. Both offset and misalignment can be measured either absolutely or relatively to the dimensions of the component or receptor. Offset and misalignment are sometimes also measured by the distance between corresponding edges of component and receptor, respectively at the beginning and at the end of the process. The (alignment) yield of capillary self-alignment indicates the fraction of components that end up being correctly positioned after process completion on a selected number of receptors (i.e., 9 correctly aligned components out

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of 10 delivered onto 10 receptors corresponds to a yield of 90%). Successful alignment is usually defined by compliance with targeted misalignment margins. The margins are in turn specifically determined by the applications at hand. Beside alignment yield, functional yield is relevant in the assembly of systems whereby the mechanical connection needs to provide also an electrical path between receptor and component. Functional assembly yield is altogether a more restrictive success metric, since typically only a subset of the correctly aligned components can also sustain electrical functionality. For instance, in the packaging of advanced integrated circuits onto back-end-of-line (BEOL)-processed substrates, sub-micrometric post-alignment registration accuracy is being targeted to enforce working electrical interconnections through solder microbumps.18 Finally, the overall efficiency of the process is measured by its throughput, which quantifies yield per unit of time. Sections 3, 4 and 6 and Mastrangeli et al.21 provide an overview of major failure modes in capillary self-alignment and fluidic self-assembly.

Wetting and Contact Lines A drop (Fig. 2a) is a finite liquid phase delimited by an interface which is the seat of the surface tension phenomenon.22 The excess surface energy of the interface is its surface tension γ. When a liquid drop is in contact with a solid substrate, two additional interfaces besides the liquid–gas (LG) are identified: the liquid–solid (LS) interface and the solid–gas (SG) interface elsewhere (Fig. 2b). The intersection between the surfaces tangent respectively to the solid–liquid and the gas–liquid interfaces is called the (triple) contact line. The angle defined by both tangent planes and spanning the liquid phase is called the contact angle θ. The contact angle governs the capillary forces exerted by the liquid on the solid (see Section 2.4). Usually, the tangent to the solid–gas interface coincides with the tangent to the solid– liquid interface. There exist however situations where this is not the case, for instance when the contact line wets a sharp solid edge (see Fig. 2c and Section 3.1). When the droplet makes contact with separated solid surfaces, it is usually referred to as a meniscus or liquid bridge (Fig. 2d).4,23

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Figure 2: (a) A suspended liquid drop (without support); (b) a liquid drop wetting a rigid substrate without solid edge, defining a triple contact line and a static contact angle; (c) a liquid drop wetting a rigid substrate with a solid edge; (d) a liquid bridge, i.e. a liquid drop bridging two separated solid surfaces. (e) Moving contact line with constant prescribed (advancing) contact angle, versus (f) pinning of the contact line at a fixed position imposed by a solid edge. The compression of the liquid bridges is in both cases consequent to the reduction of the gap between the parallel bounding surfaces.

The contact line is in principle free to move across a substrate with chemically homogeneous and topographically smooth surface (Fig. 2e). At equilibrium, in proximity of the contact line the meniscus will eventually assume on a sufficiently rigid substrate a shape set by the equilibrium contact angle, which, although ill-defined, is typically described by the Young–Dupré equation cos θ = (γSG − γLS)/γ. The equation captures the balance of the tension components parallel to the substrate. A low θ signals high chemical affinity of the liquid to the substrate, and perfect wetting is achieved for θ → 0. Over rough and/or chemically heterogeneous surfaces, a range of apparent contact angles can instead be observed at a given contact line position (Fig. 2f). Contact line movement will then only occur when the apparent contact angle is either larger than an upper threshold value, called the advancing contact angle θa, or smaller than a lower threshold value, called the receding contact angle θr. The angular sector [θrθa] between both threshold values quantifies the contact angle hysteresis. When the contact angle lies within such interval, the contact line is said to be pinned. Contact line pinning, whose microscopic origin is attributed either to localized defects or adhesion hysteresis,24 can be further reinforced by the presence

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of sharp edges on the substrates (Fig. 2c). The increases in contact angle hysteresis induced in this case by slope discontinuity can be exploited for topographical liquid confinement, as described in Section 3.1. For a geometrical description of meniscus shape, the knowledge of the contact line position is needed when the contact line is pinned (Dirichlet boundary condition), while the knowledge of the contact angle must be used when the contact line is moving or its position is unconstrained (Neumann boundary condition).23 The dynamic contact angle is also dependent on the velocity of the contact line.25 The contact line can switch from being pinned to mobile during a single experiment according to changing conditions. For instance, in the early instants after deposition on a homogeneous solid surface, a droplet relaxes and the liquid spreads over the substrate until its motion stops upon reaching an equilibrium configuration or a boundary. Similarly, when a component is set on top of a confined liquid droplet, either by release or by pre-contact,26 wetting of the bottom surface of the component starts upon contact with the liquid interface, and the liquid stops spreading when the meniscus reaches the edges of the component (Fig. 1, bottom row). In the following subsections, we show how analytical descriptions of capillary forces can be derived from surface energy- and force-based models for the simple case where the liquid meniscus completely wets the receptor and the component. In Section 3.2 this assumption will be relaxed.

Surface Energy Models The mechanics of capillary self-alignment can be understood as the consequence of the minimization of the system’s excess free energy. Specifically, the surface energy of the liquid meniscus is minimized when its interfacial area is minimized. In a proper design of the system, this condition is satisfied when the edges of the component are aligned to those of the receptor. Assuming other energy components to be negligible (e.g., gravitational energy, for Bond number Bo ≪ 1), the excess free energy is represented by the interfacial energy E of the system

E = γA (1) where γ is the surface tension and A is the total interfacial area of the liquid. Assuming to know the equilibrium shape of the liquid meniscus for each = [x0y0z0] of the component, the restoring force offset calculated from the energy function E = E( ) by

can be

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(2) Considering the energies of all interfaces, eqn (1) is expressed as: E = γALG + γLCALC + γLRALR (3) where γ** and A** represent respectively the interfacial energies and areas among liquid (L), gas (G), component (C) and receptor (R) (see Fig. 1b). In case of e.g. fluidic self-assembly, the gas phase can be replaced by the immiscible liquid hosting the process.27–29 Under the common and simplifying assumptions of perfect wetting (Section 2.2), which implies full and constant liquid coverage of the liquid–solid interfacial areas (ALC and ALR), and absence of surface tension gradients, eqn (2) reduces to: (4) Eqn (4) can be used for estimating self-centering forces for simplified geometries of liquid menisci, such as given by circular30,31 or rectangular2,32,33 components on shape-matching receptors under purely lateral translation (Fig. 3). By approximating the shape of the corresponding menisci as a sheared cylinder and parallelepiped, respectively, the energy can be written as: (5) (6) For small displacements (x ≪ h), the lateral force Fx = −dE/dx can be approximated as: (7) (8) that is, the lateral restoring force is linearly proportional to the displacement from equilibrium, as in an Hookean spring, and to (half of) the perimeter of the component/receptor. Following this approach, Berthier et al. generalized eqn (5)–(8) to the case of arbitrary polygonal components.34 Gao and Zhou proposed a unified semi-analytical model that describes small deformations of liquid menisci through plate series.35 It should be emphasized that these

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are only coarse approximations, since real menisci are not ideal cylinders nor parallelepipeds. More accurately, the energy E has to be solved by finding the actual shape of the liquid meniscus under the given boundary conditions. In the most general cases, the solution can only be found computationally. This requires modelling the meniscus by digital tools, such as multiphysics software or Surface Evolver (ref. 2 and 8 provide extensive descriptions and examples). Numerical simulations additionally allow the characterization of the fluid geometries for all roto-translational degrees of freedom.4,32,36

Figure 3: Simplified meniscus geometries for the analytical derivation of the capillary force from the surface energy: case of a circular (a) and of a rectangular component (b) over shape-matching receptors.

Capillary Force Models In the previous section, the capillary self-alignment process was analysed through energy minimization, and centering forces were derived through the energy gradient. The centering capillary forces can alternatively and equivalently be calculated directly from the composition of their two contributions: the tension force and the Laplace force. Indeed two distinct concepts merge under the denomination of capillary forces: the tensile, tension or surface tension force FT, and the pressure, Laplace or simply capillary force FL. The reader must bear in mind these two different physical effects, adding to one another to constitute what we will call the capillary force in the following. The first contribution FT is due to the tensile force exerted by the liquid meniscus on the solid and directed along the tangent to the liquid–gas interface. The surface tension force, always tensile, is quantified by the integral of the surface tension along the triple contact line (see Fig. 4 for symbols definition):

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(9) According to recent studies by Marchand et al.,37 this is to be considered an imprecise physical description of the capillary effect of a liquid on a compliant solid. We will however use the traditional description in the following, since both theories agree with one another for rigid solids.

Figure 4: Free body diagram of the component showing the tensile and pressure components of the capillary force. The tensile force is computed as the line integral of the force element d T = γdl t(see inset). The Laplace force is computed as the surface integral of the force element d L = 2HγdS Nwhere, according to the depicted sketch, Δp = 2Hγ is negative. Unit vectors are defined in the inset..

The second effect is associated with the so-called Laplace pressure drop Δp, induced by the curvature of the meniscus.38 The Laplace pressure drop is given by the well-known Young–Laplace equation: pin − pout = Δp = 2Hγ

(10)

where 2H is the total curvature, equal to the sum of the orthogonal curvatures 1/ρ1 and 1/ρ2.2,22 This leads to a pressure-related force FL equal to: (11)

These concepts are exemplified in the snapshot shown in Fig. 5. In this case the surface tension force is estimated to equal the perimeter of the square chip (800 μm) times the surface tension of water (72 mN m−1 at room temperature) projected along the symmetry axis by an angle of 10°, i.e. FT =

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59 μN along the diagonal defined by . Estimating the radii of curvature to be ρ1 = 105 μm and ρ2 = 272 μm leads to a curvature 2H = 13 200 m−1, and a Laplace pressure Δp ≅ 950 Pa. This provides a Laplace force FL = 38 μN along the diagonal

force is FT + FL = 21 μN along the diagonal

. The resulting capillary .

Figure 5: Force analysis on the capillary self-alignment study described in ref. 39. (a) Estimation of surface tension force, (b) estimation of Laplace pressure effects (adapted from ref. 39, © 2012, American Institute of Physics).

Model Equivalence The modelling methods presented in Section 2.3 (energetic) and 2.4 (sum of Laplace and tension force components) are demonstrably equivalent. As a quick illustration, let us consider the configuration of Fig. 3b. The Laplace pressure acts perpendicular to the bottom surface of the component and does not contribute to the lateral restoring force along x. The tension force only contributes to the total capillary force along x, and can be expressed as: rect

≈ 2cγ cos α

x

(12)

for x ≪ h, leading to rect ≈ where −2cγx/h x as in eqn (8). A more detailed analysis of the equivalence of energy- and force-based models is reported in ref. 40.

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LIQUID CONFINEMENT Liquid confinement amounts to imposing a barrier to the spreading of a liquid on a substrate. To be localized on the target substrate, a liquid droplet must be confined along its entire perimeter by the edges of the receptor. Without confinement of the liquid meniscus, the final position of the component is underspecified and not predictable. Only in presence of liquid confinement can the shape of the receptor conveniently match the profile of the component.

Receptor Types and Fabrication Liquid confinement on a solid substrate can be enforced in a surface-chemical way, in a geometrical way, or in a combination thereof.26,41 In all cases, the mechanism underlying liquid confinement is the pinning of the contact line and the associated extension of the contact angle hysteresis along the pinned perimeter of the droplet (Section 2.2). The surface-chemical method (Fig. 6a) exploits the wetting contrast between the interior surface of a receptor (i.e., the receptor proper) and the exterior surface, which can be intended e.g. to space arrays of receptors apart according to a defined pitch.42The receptor surface needs to be significantly more wettable by the liquid than the spacer surface. The difference between the liquid contact angles on the adjacent surfaces provides a rough quantification of the wetting contrast. According to this metric, a difference larger than 60° is sufficient for robust chemical confinement.42–44Once on the edge of the receptor, the contact line remains pinned for all values of the edge angle comprised within the canthotaxis sector, delimited respectively by the receding contact angle on the wetting side and by the advancing contact angle on the non-wetting side.8 Hydrophilic (Fig. 7a and b) or oleophilic receptors (Fig. 7c–e) are specifically implemented to be highly wettable by water or hydrocarbons, respectively. Chemical confinement is effective also for liquids of low surface tension in air.45

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Figure 6: Liquid confinement within receptors. (a) Surface-chemical confinement, exploiting the wetting contrast between adjacent surfaces. (b) Topographical confinement, using the geometry of a solid edge of aperture φ. (c) Implementations of topographical liquid confinement.

Figure 7: Microfabricated receptors for surface chemical (a–e) and topographical (f–h) liquid confinement. (a) Top view and (b) side view of a hydrophilic silicon dioxide receptor surrounded by hydrophobic fluorocarbontreated spacer. Water droplets wet only the hydrophilic receptor (adapted from ref. 54, © 2011, Institute of Physics). (c) Oleophilic Au-coated receptors

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surrounded by oleophobic porous spacer functionalized with a fluorinated selfassembled monolayer (adapted from ref. 45, © 2011, American Institute of Physics). (d) Side view of a droplet of oil-like adhesive on Au, and (e) droplet of the same adhesive on the background (adapted from ref. 45, © 2011, American Institute of Physics). (f) Topographical confinement of a water droplet within a laser-scribed receptor on plastic substrate (adapted from ref. 55, courtesy of G. Arutinov). (g) SEM micrograph of a receptor topographically defined by undercut edges (reprinted with permission from ref. 56, © 2013, IEEE). (h) An oil-like adhesive droplet confined on a receptor defined by an undercut edge, showing ∼180° apparent contact angle (adapted with permission from ref. 49, © 2013, Wiley).

The geometrical or topographical method makes use of sharp solid edges of the substrate to define the shape of the receptor46 (Fig. 6b). In this case, the edge angle of the liquid needs to increase up to the advancing contact angle value with respect to the plane tangent to the outer side of the receptor before the contact line can further advance, causing liquid overflow out of the site. Such extension of the contact angle hysteresis is captured by the so-called Gibbs’ criterion, which for a chemically-homogeneous solid surface defines the canthotaxis sector as [θr, θa + π − φ], where φ is angle spanned by the solid edge.47 The most common geometries of topographical receptors include trench,48 mesa and undercut types (Fig. 6c and 7f, g). Notably, by using receptors with solid undercuts, it is possible to confine low surface tension liquids in air, and obtain apparent contact angles in excess of 180° (Fig. 7h).49 Sparse arrays of micropillars with a doubly-reentrant nanomachined tip geometry were recently demonstrated to turn every solid surface, including highly wetting ones, superrepellent to all liquids.50 An issue for mesa-type receptors of micrometric size is the required step height difference, which should be of the order of 1 μm.51 Receptors exploiting a hydrophilic/hydrophobic wetting contrast are relatively easy to fabricate using photolithography. Conversely, the more reliable hydrophilic/ superhydrophobic contrast is challenging to enforce due to the surface nanoroughness required in the spacer. Recently, electron beam was employed to achieve patterned wetting and create μm-sized superhydrophilic patterns on superhydrophobic substrates.52 Hydrophobic sidewalls surrounding hydrophilic mesa-type receptors were also shown to improve liquid confinement and benefit the capillary self-alignment performance.26,41 Similar observations were shown to apply to the surface conditioning of components.53

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Partial Wettability In presence of perfect wetting (θ → 0, Section 2.2), the contact lines of the liquid meniscus are pinned along the edges of both receptors and components. The perfect wetting condition represents a useful model simplification and a good approximation for common applications, such as e.g. solder bumps over oxide-free metal pads or water on highly hydrophilic receptors (see previous section). However, perfect wetting does not capture the most general scenario.57 Upon dispensing, a large advancing contact angle may prevent the liquid to wet the complete receptor. Capillary self-alignment was nonetheless demonstrated also on partially wetting or hydrophobic receptors, such as e.g. with advancing and receding water contact angles of 118° and 69°, respectively.39 Wetting of the receptors was in this case forced by using excessive amounts of liquid or by applying pressure on the component. Capillary self-alignment across relatively large component offsets is another significant instance where partial wetting affects the performance of the process in terms of accuracy and yield.48 Partial wetting makes the liquid/solid interfacial areas time-variant during transient component motion26,57 and, as expected from eqn (3), this affects lateral restoring forces.58 A recent piecewise model of lateral restoring capillary forces can justify the experimental observations by accounting for partial liquid wetting.33 According to what was mentioned in Section 2.2, in this case the boundary conditions must be adapted to the offset of the component to reproduce the local meniscus geometry in proximity of the contact lines. In the model, progressively larger lateral deformations of the liquid meniscus, corresponding to progressively larger component offsets, are accommodated by switching the boundary conditions from pinned contact line to prescribed receding contact angles.33

INFLUENCE OF SHAPE AND SIZE Capillary self-alignment typically requires the shape of the component to be commensurate with that of the receptor. The general design guideline is that, for a given initial component offset, the liquid meniscus constrained by receptor and component (or a part of it) possesses sufficiently higher potential energy compared to its equilibrium shape so that its relaxation can sustain the self-aligning motion, as discussed in Section 2. Most reported instances of the process target perfect final alignment between the components and

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the receptor. Accordingly, components and receptors have shapes and sizes matching up to the degree allowed by fabrication process,29,48,54,59–61 which is often in the order of fractions of μm. In some applications, the receptor is fractioned into several separated parts, due to the requirements of electrical functionality, structure of the device, or to remove excessive amount of liquid after alignment.62,63 The working principle for such single-component to multi-receptor embodiment is analogous to the standard single-component to single-receptor self-alignment configuration, except for the mechanism of meniscus formation, which may slightly differ due to the segmented shape of the receptor.62

Capillary Length and Liquid Volume The volume of the liquid meniscus is an important parameter for achieving high yields from the self-alignment process. To discuss the appropriate liquid volume for self-alignment, the concept of capillary length is useful, where γ is the surface tension of the droplet, ρ is the density of the liquid and g is the gravitational constant. For a water droplet in air at room temperature, the capillary length is about 2.7 mm. Capillary length can be used to quickly check the dominating force in self-alignment regarding the different size of patterns. For example, when the components and receptors are smaller than the capillary length, and so is the gap, it is safe to say that the surface tension will dominate the self-alignment process. Similarly, by replacing the gravitational constant g by an acceleration, an equivalent cutoff length can be defined, which is relevant for high-acceleration pick-andplace machines.6 However, LC alone cannot be used to determine whether selfalignment will succeed or not. Other important parameters are volume and corresponding thickness of the meniscus between component and receptor, surface wetting, the weight of the component and the possibility of friction between solid surfaces. Self-alignment of rather thin, 20 mmsized components (i.e. an order of magnitude larger than LC of water) was reported with good success rates.20,48,64,65 Even below the capillary length and in absence of liquid overflow, an excessive liquid volume can cause a tilted and misaligned component pose, corresponding to a locally stable meniscus geometry.66,67Conversely, at the small droplet limit there is not enough liquid to wet both bounding surfaces, the component can touch the substrate and dry friction can prevent alignment completion. For maximizing yield of self-alignment, there may be an optimum liquid volume value between

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these extremes. To the knowledge of the authors, only a few design rules for selecting the droplet volume have been provided.67 For the specific case of 300 μm × 300 μm cuboidal components assembled over square shaped receptors, the optimal droplet volume maximizing yield was experimentally found between 1.8 and 2.7 nL,60whereas a gap of 125 μm was suggested to optimize the performance in the alignment of millimetre-sized foil dies.64 The addition of wetting bands around the edges or corners of the receptors was suggested to help avoid tilting of the components.66

Shape Matching Self-alignment does not require the shape of the components and of the receptors to strictly match each other. In early studies,28 the receptors were designed with shapes supposed to allow robust correction of initial angular offsets, such as half-discs and commas. The patterns could not exclude unintended component placements corresponding to local energy minima. Analytical studies on the design of component and receptor shapes were pursued by Böhringer et al.68,69 They developed a simple geometrical model to predict the energy landscape for the self-alignment of components and receptors based on the convolution of their shapes. The model is valid for arbitrary geometries and for large biases compared to the gaps. The model recovers the intuition according to which the energy of the system is smallest for the largest overlap between component and receptor, and is helpful to predict and avoid configurations associated with local energy minima.69 For the widely used cuboid-shaped components and square or rectangular receptors, capillary self-alignment faces several application scenarios depending on the feeding methods (Section 6.1). Particularly, if the shapes and sizes of the component and the receptor are significantly mismatched, besides local energy minima there can exist multiple relative positions which cannot be distinguished by their energy. In presence of such degeneracy, the final position is contingent to the details of feeding and can be hardly predicted with precision. Fig. 8 shows general scenarios wherein the size of the components and of the receptors are significantly different and their final relative position is underconstrained. Cases A1–B3 represent a component placed onto a larger receptor. When released partially outside the receptor, the component can still marginally align to the receptor, provided that the motion is overdamped. This can lead to partial self-alignment (A1), corner self-alignment (A2) or edge self-alignment (A3).70,71 The component may conversely overshoot

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past the receptor edge and pull the contact line with it, away from the edge, when small liquid viscosity and/or large liquid volumes induce underdamped motion. Insufficiently large offsets may also not be conducive to the alignment conditions A1–A3. In these cases, the component may end up in either of the B1–B3configurations (B2 being equivalent to A3). When the weight of the component is insignificant compared to surface tension of the meniscus, the relative sizes of components and receptors can also be inverted, so that for instance the configurations C1–C3 can be mapped onto A1–A3. It is interesting to note that in all cases, except for B3, the in-plane orientation of the component will be corrected during the self-alignment.

Figure 8: Scenarios for capillary self-alignment where the sizes of component and receptor are not commensurate. (A0) Default: component and receptor with matching shapes; (A1) matching of a single lateral size, leading to partial selfalignment; (A2) component matching the shape of a corner of the receptor, leading to corner self-alignment; (A3) component matching the shape of an edge of the receptor, leading to edge self-alignment. (B1–B3) Components released from undesired offsets, leading to poor self-alignment or even no alignment. (C1–C3) Reciprocal cases to A1–A3, leading to similar results only in absence of component tilting and of friction between dry surfaces.

Capillary self-alignment can also be used to align components with more sophisticated geometries than the primary shapes previously analysed.72 Millimeter-sized MEMS force sensors were picked up and precisely aligned to a pre-patterned printed circuit board (PCB).73 The capillary force of molten solder bumps between matching pads on both board and

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sensors was exploited, and three devices were assembled simultaneously. A capillary approach was also applied to the fluidic assembly of a GaAsbased microcantilever spin injector,70 microcantilever for atomic force microscopy,71 and of millimeter-sized components with inner cavities.63 In this case the hydrophilic receptors, coated with a thin layer of water-diluted hydrofluoric acid, matched only the perimeter of the components. The receptors contained additional microchannels for liquid evacuation, so that an interior area free of liquid could safely be maintained to host additional structures. This solution is suitable for sealed packaging of e.g. sensors and MEMS devices. When the size of the component and the receptor are commensurate, the self-alignment can allow accurate positioning despite mismatch in shape or edge profile (Fig. 9). Defects inside the receptor or along the edge of the receptor bear relatively little influence to the alignment.74 Even when the edge of the receptor is poorly defined, due to e.g. low quality manufacturing, the self-alignment will still lead to rather good accuracy. For example, for receptors of size 200 μm × 200 μm with edge jaggedness of ±4 μm, the lateral and orientational self-alignment accuracy of smooth matching components was ±1 μm and ±0.6° respectively, which is significantly better than the definition of the pattern. The reason for this is that the energy minimization of the self-alignment process arises from a weighted average where the influences of opposite edge distortions may cancel each other out, leading to a modest magnitude of the total error.

Figure 9: The alignment accuracy of 200 μm × 200 μm × 50 μm components on 200 μm × 200 μm receptor with different edge profiles: linear, jaggedness of ±2 μm, ±4 μm and ±8 μm (adapted from ref. 74, © 2014, IEEE).

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Size of Components Capillary self-alignment can be applied to a large range of component sizes, from tens of μm to tens of mm, and to components with length to thickness ratio up to 100.

Large components Capillary self-alignment works best when surface tension dominates. When the characteristic size of the component is larger than a few mm, gravitational and inertial effects may become comparable (Section 4.1). Nevertheless, self-alignment of relatively large dies has been successfully demonstrated. In this case, the thickness of the dies is usually below 1 mm, and the height of the liquid meniscus measures few hundreds μm, such that capillary forces still prevail. Assembly of arrays of 20 mm dies on carrier substrates through a combination of parallel vacuum gripping and capillary self-alignment was reported.65 The capillary self-alignment of 10 to 20 mm-sized plastic foil dies fed by pick-and-place onto silicon-based64 and laser-scribed plastic substrates48 was also demonstrated, the latter being also suitable for autonomous deployment onto moving webs.75 Centimeter-sized components floating at the air/water interface were assembled onto matching hydrophilic receptors of hydrophobic substrates.76 The substrate was pulled at an angle through the fluid interface, and unique pre-orientation of the components could also be enforced by using a magnetic template.77 Capillary self-alignment is also simplifying 3D microelectronic integration. For instance, several layers of 5 mm × 5 mm × 50 μm chips with 5 μm-diameter through-silicon VIAs (TSVs) were self-aligned and stacked onto BEOL-processed silicon substrates.19

Ultra-small components 20–60 μm-sized semiconductor components were assembled on a vertically pulled substrate at a silicon oil/water/substrate triple interface.59,78 The surface chemical treatment of the components and the chosen immiscible fluid couple made the trapping of the chiplets at the interface energetically favourable. The authors reported the largest self-assembly throughput to date with 62 500 chips assembled in 45 s, equivalent to 1388 s−1. The process requires significant preparation steps for each assembly run, and repeated crossing of the substrate surface by the liquid interface carrying the chiplets

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is needed to achieve the claimed high surface coverage of the receptors. The reported accuracy and assembly yield were 0.9 μm and 98%, respectively.

Ultra-thin components Capillary self-alignment can be particularly suitable to handle very thin and fragile components, since gravity is small for the otherwise significant size of the chips. 12 ultra-thin 10 μm chips were successfully stacked and precisely aligned on top of each other using robotic pick-and-place and capillary selfalignment.79 Capillary forces were also combined with magnetic alignment, mediated by magnetically-susceptible metal interconnects, for the templated integration of 20 μm-thick components onto polymer foils.80

DYNAMICS Inertial and Viscous Effects A no-flow situation was implicitly considered in the previous sections, with the liquid meniscus at mechanical equilibrium. To describe the physics of the process in presence of liquid flow in the meniscus, the associated Navier– Stokes equations must be solved. The viscous effects induced by the flow can then be compared to the capillary effects through the capillary number , while the relative magnitude of inertial over viscous effects is considered in the Reynolds number , with L the characteristic dimension of the system (typically, the component sidelength). The comparison of capillary, viscous and inertial effects was studied by Valsamis et al. in the case of axial motion of liquid bridges,81 illustrating the existence of three distinct regimes: a capillary regime for small Ca and small Re, a viscous regime for large Ca and small Re, and an inertial regime otherwise. It remains unknown whether this description is valid also for lateral motion. To go beyond pure scaling laws, it is necessary to consider the set of Newton and Navier–Stokes equations describing the alignment process. For the simplified cylinder and parallelepiped geometries of Fig. 3, where only the lateral displacement of the component during the self-alignment motion is considered, it is possible to develop a simplified analytical approximation of the viscous force82 by assuming Couette flow between the component and the receptor. Assuming a linear velocity profile, the viscous shear stress on

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the component can be written as force is thus equal to:

, and the viscous

where μ is the dynamic viscosity, ALC = wl and w, l, x and h are respectively the component width, length, offset and the gap. For the parallelepiped case of Fig. 3b, this can be combined with Newton›s law and the linear approximation of the surface-tension force from eqn (8) to describe a damped second-order dynamics of the capillary self-alignment process: (13) From this model, one can estimate the natural frequency ω and the damping coefficient ξ of the dynamical system: (14) (15) These results can thus be interpreted as usual for a second-order system dynamics. The linear velocity assumption used above only holds when the flow velocity has the time to develop a linear profile, i.e. after a timescale of about h2/ν, ν being the kinematic viscosity. If the restoring capillary force triggers oscillations with a shorter period, which can be the case for large gaps or low viscosities, the coupled equations must be solved simultaneously.30,83 The results are represented in Fig. 10, where is a characteristic time of the oscillations of the component mass m subjected to a capillary is stiffness k (e.g. for rectangular components, a non-dimensional viscosity (corresponding to 4 times the ratio of the is a non-dimensional characteristic time τ and the timescale h2/ν), mass (component mass scaled by the liquid mass). The map then gives the

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non-dimensional characteristic time

247

associated to the second-order

when eqn (13) system defined above, where the characteristic time holds. Table 1 lists pertinent values of Ca and Re extracted from recent literature.

Figure 10: Non-dimensional characteristic time τd as a function of non. The stripped lines indicate the analytidimensional viscosity α and mass cal solutions to eqn (12) of the linear second-order system, valid for α > 1. , the solid lines are obtained from numerical simulations30 For smaller and deviate from the analytical expression [reproduced from ref. 2, © Springer, 2013]. Table 1: Capillary and Reynolds numbers from selected recent works Reference Lu and Bailey (2005)83

Ca 4.62 × 10−5 0.012

Re 0.933 0.055

Lambert et al. (2010)30 Silicone oil Lambert et al. (2010)30 Silicone oil

0.158

0.010

Arutinov et al. (2014)20 Water Water Dubey et al. (2016)84

1.39 × 10

0.013

1.22 × 10

0.674

Lambert et al. (2010)30

Liquid Solder paste Glue Dymax

75 404

0.002 −6 −4

The in-plane modes of oscillation of a millimeter-sized self-aligning component were recently tracked simultaneously and with high spatial

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and temporal resolution by means of a high-speed camera (Fig. 1, bottom row).20 The study evidenced the possibility of both uncoupled and coupled motion along the degrees of freedom of the fluid joint depending on the initial conditions of component release. As noted above, increased viscosity and lower surface tension slow down the capillary self-alignment process, making it easier to visualize using ordinary frame rate cameras.

The full dynamics of chip alignment is based on the coupling of the 6 degrees-of-freedom of the component (3 translations and 3 rotations) with the flow dynamics of the meniscus. This amounts to coupling Newton’s law for the component with Navier–Stokes equations for the fluid. Analytical estimates for the axial motion81,82 and, under strong assumptions, for the lateral motion82 were reported. This is however not representative for the transient period typical of the oscillatory motion of a self-aligning component, as observed by Arutinov et al.20 Other solutions were proposed by Lu and Bailey83 and Lambert et al.30 relying on the assumption of one-dimensional liquid flow, i.e., of liquid flowing only parallel to the fluid film (v‖ ≠ 0, v⊥ = 0). None of these approaches however manages to describe very accurately the dynamics of capillary self-alignment of the component recorded experimentally.20 Viscosity and Surface Tension of Common Liquids The previous analyses show that several physical and chemical parameters of the liquid menisci influence the self-alignment process: surface tension (typically ranging from some tens of mN m−1 for water-based solutions, adhesives and oils up to several hundreds of mN m−1 for molten metallic alloys), dynamic viscosity (ranging over several orders of magnitude) and density. For the most commonly used liquids we specifically report in Fig. 11 the values of surface tension and viscosity, which are explicitly involved in the analytical models (Section 2.3). It is apparent that the liquids can be grouped into two main clusters at opposite ends of the ranges, and that water sits roughly in between. It is also apparent that the viscosity varies up to 4 orders of magnitude between liquids, while the surface tension varies only over 1 order of magnitude. As seen in the previous section, the capillary number and the characteristic time constant are set by the ratio of viscosity and surface tension. Hence, since there is greater freedom in choosing liquid viscosity, to tune the self-alignment time it is usually preferred to change the viscosity of the liquid, all other parameters being equal.

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Figure 11: Dynamic viscosity versus surface tension for most used liquids in capillary self-alignment. Typical orders of magnitude for adhesives and solders (rectangular boxes) are taken from ref. 85. Values for Dymax glue, silicone oil and molten alloy Indalloy can be found in ref. 30, and for epoxy adhesive in ref. 86. Liquid density values (needed to compute e.g. the Reynolds number, see Table 1) are indicated in the legend.

Experimental Self-alignment Dynamics As earlier mentioned, most of the works in capillary self-alignment have considered perfect wetting of liquids on both receptors and components (e.g., oxide-free molten solders on metal pads, water on oxidized surfaces), and excluded contact line unpinning and partial wetting during the process. In presence of contact line pinning, contact angles affect the self-alignment process only for what concerns liquid confinement, since vanishing contact angles and small wetting contrast tend to make the confinement unreliable and not robust48 (as discussed in Section 3.2). Conversely, upon delivery of the component on top of the confined droplet, the contact lines are generally not pinned, and the meniscus transiently spreads to reach the edges of the component (Fig. 1, bottom row). The dynamics of self-alignment is then significantly affected by liquid spreading, and thus by the contact angles.58 Higher wettability (i.e., lower contact angles) was shown to induce faster aligning motion of the components,20,58 consistently with the predictions of energy-based models.33However, self-alignment can still be achieved through the relatively slow spreading of a liquid medium of high viscosity and low surface tension, such as resin materials86,87 or adhesives.29,45Fig. 12 illustrates the capillary self-alignment between two shape-matching and vertically stacked 300 μm × 300 μm SU-8 chips bridged by an adhesive of

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high viscosity curable by ultra-violet (UV) radiation. The process reached successful completion about 4 minutes after release of the top chip.

Figure 12: Snapshots from the recording of the capillary self-alignment of two stacked and shape-matching 300 μm × 300 μm SU-8 chips mediated by an interposed highly viscous UV adhesive.

PROCESS INTEGRATION The self-alignment process includes several essential process steps: component feeding, liquid deposition, self-alignment and post-processing. The order of component feeding and liquid deposition is determined by the design of the whole process.

Liquids Liquid deposition and choice To achieve surface tension-driven self-alignment, an appropriate volume of liquid must be introduced at proper time during the process. Liquids are conveniently introduced onto the receptors (see Section 3) and, especially in presence of very small volumes of volatile liquids, the component should be introduced shortly after. Liquid deposition methods can be divided into four main categories: (a) non-contact, single-droplet methods; (b) contact, single-droplet methods; (c) parallel stochastic methods and (d) parallel deterministic methods. Fig. 13 illustrate the working principle of the different liquid deposition methods, and Table 2 summarizes their main features.

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Figure 13: Common methods of liquid deposition: (a) non-contact, single droplet; (b) contact, single droplet; and (c) parallel stochastic; (d) parallel deterministic. Table 2: Common methods for liquid dispensing Single droplet, non-contact

Single droplet, contact

Parallel stochastic

Parallel deterministic

Technologies

Droplet dispenser Time pressure valve, pump

Condensation, mist generator, spray

Dip coating

Minimum dispensed liquid volume

∼pL

∼pL

∼fL

Down to fL

Process dependent, e.g. 1.5 nL mm−2 s−1

Dependent on pulling speed

Liquid viscosity

Aqueous to high viscosity (e.g. 500 000 mPa s)

Most liquids (instrument dependent)

Aqueous (technology dependent)

Aqueous (technology dependent)

Positioning

1D or 2D

2D or 3D

2D

Vertical pulling

Comments

Dispenser-to-substrate distance: few- to tens of mm

Contact influences positioning speed

Dispensing covers the whole substrate

Liquid should de-wet from the substrate

References

88–90

91

54 and 92

42, 59 and 93

Liquid deposition speed

1000 s droplets per s

Continuous flow

In non-contact, single-droplet methods a droplet is shot from a nozzle to the receptor, similarly to inkjet printing (Fig. 13a). Numerous receptors can be addressed efficiently and in sequence with the help of a motorized

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positioning system when the dispenser is shooting at a high speed. The resulting single droplet volume can measure down to tens of picoliters.89,90 Such non-contact dispensers are suitable for water-based and low viscosity fluids (e.g. 5 mPa,90 20 mPa89). More viscous liquids can also be handled, provided their viscosity is reduced (e.g. by heating) below a dispenserdependent threshold value.88 During operation, the tip of the dispenser is kept a few millimeter away from the receptors. A distance of tens of millimeters is possible with reduced reliability. Acoustic liquid handling94 is an alternative technique, in which high-frequency acoustic signals are focused on the surface of a fluid of interest to eject sequentially and directionally droplets of nL to pL volumes with high accuracy and precision.95 Contact, single-droplet methods (Fig. 13b) include a variety of dispensers and techniques, from traditional time-pressure dispensers and pump dispensers to nano-contact printing. In all these methods, a single droplet is formed when the nozzle of a needle comes into contact with the receptor, followed by liquid injection, pinch-off and removal of the needle. Nanoliter droplets are typically achievable with commercial dispensers. With more precise control of the contact time profile, smaller volumes are also achievable. Similarly to non-contact methods, the addressing of multiple receptors require proper positioning systems. Contact dispensers can dispense a wide range of liquids, from aqueous to very viscous. Parallel stochastic methods (Fig. 13c) provide an efficient way to deposit at once liquids to a large amount of receptors or component-receptor pairs. In parallel stochastic methods, liquid is delivered randomly to the substrate e.g. by condensation from vapor phase or by water mist droplets impacting on the surface. Compared to the single droplet methods, the parallel stochastic methods can produce a huge amount of individual droplets of down to femtoliter volumes, which makes it highly efficient. The parallel stochastic method is most suitable for water disposition,54,92,96 even though other liquids can also be handled. As the deposition takes place on the full substrate, droplets on undesired regions of the substrate (e.g., outside the receptors) should be removed by post-processing (e.g., vaporization) after the self-alignment step, unless their existence is compatible with the subsequent processing steps. Parallel deterministic methods, such as dip-coating (Fig. 13d), can deposit liquid to a large amount of receptors.42,59 In dip-coating, the substrate patterned with receptors is dipped into the liquid to be coated and then redrawn from the liquid.44,93 Depending on the receptor types,

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either chemically of topographically defined (see Section 3.1), the coating dynamics and particularly the failure of the liquid bridges can change significantly.44 In the case of chemical patterning, the receptors should be hydrophilic, while the substrate should be sufficiently hydrophobic such that only the receptor will be coated and the substrate will be dewetted by the liquid. Compared to e.g. lines, the coating of closed chemically-patterned patches, like the receptors, involves capillary break-up effects that are intrinsically time-dependent.43 The thickness of the coated liquid depends on the pulling speed and insertion angle of the substrate, the surface tension, viscosity and density of the liquid, the surface energy of the receptors and the substrate (i.e., the wetting contrast), and the shape, size and orientation of the receptors.42,43,93 Coating thickness down to 1 μm (equivalent to 1 nL mm−2)93 and sessile droplets of femtoliter volumes44 can be obtained by dipcoating. Additionally, low melting point solders are typically deposited on receptors by this technique.59Selective liquid deposition onto hydrophilic receptors can also be carried out using droplets sliding over hydrophobic substrates.97 Many parameters guide the liquid choice, including the ability to wet the chosen surfaces, electrical or thermal conductivity, optical transparency, the ability to make permanent bonding through phase change (e.g. by freezing, optically- or thermally-initiated cross-linking), and processing temperatures. The three main types of liquids that have been applied in capillary selfalignment are solders, water and other solvents, and adhesives. Their specific properties are discussed in the following sections.

Solders Capillary self-alignment mediated by solder droplets in air was the first extensively studied self-alignment process, in the context of flip-chip98–101 and surface-mount reflow soldering.102,103 The self-centering force of the solder joint was clearly identified in early work103 focusing on soldering defects. One of the major applications identified for solder self-alignment was for passively aligned optical interconnects.104The high accuracy requirements of optical interconnects could not be achieved using solder self-alignment alone, but capillary self-alignment was usually combined with mechanical shape matching or stoppers to achieve accuracies better than 1 μm,100,104suitable for interconnects. Variations in solder volume can cause undesired variations in chip-to-substrate gap, or tilts when multiple solder balls are used. These variations can be avoided by using mechanical spacers.105

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Molten solders characteristically have high surface tension and low viscosity (Fig. 11), and are therefore well suited for self-alignment. Additionally, solders can make permanent bond upon cooling and serve as an electrical connection (Section 5). Since solder balls are also used as electrical connectors, typically multiple solder joints are used, laid out in a row or a matrix pattern (ball grid arrays). The multipurpose role of solder joints means that solders are usually not chosen solely based on the selfalignment properties. Environmental aspects also limit the choice: leadbased solders are banned or limited in commercial use, e.g. by the Restriction of Hazardous Substances Directive (RoHS) in the EU. Examples of solders used for self-alignment include Sn/Pb (5/95100 or eutectic 63/37105), Au/Sn106 (78/22107 or eutectic 80/20108,109) and In/Pb (50/50).110 One issue with traditional solders can be the required high melting temperatures, not compatible with heat-sensitive materials and components. Low melting point solders,111,112 such as Bi/Sn/Pb (46/34/20)73 are therefore used. Another issue, often hindering solder self-alignment and interconnect reliability, is insufficient wetting or adhesion caused by solder or pad oxidation. This can be solved by using liquid flux during reflow, or by removing the oxide prior to the reflow and self-alignment in protected environments104 (see also Section 6.3).

Water and other solvents As for solvents, water droplets in air have mostly been used for capillary self-alignment. Water is a polar solvent and has a relatively high surface tension, resulting in large forces for self-alignment. Water is an attractive choice also because it is easily available, non-toxic and compatible with many materials and processes. Moreover, (super)hydrophobic surfaces can be fabricated in many ways and can be used as non-wetting areas for selfalignment substrates (Section 3.1). Water alone does not achieve a permanent bond, even after evaporation. By using silicon oxide surfaces, and adding small fractions of hydrogen fluoride (HF) into the water, components can be chemically and permanently bonded to receptors.113Alternatively, several authors have explored a twophase process,114–116 whereby the initial self-alignment is done using large hydrophilic receptors and commensurate water droplets. Metallic (e.g. Au finished) bonding pads and corresponding solder bumps are contained inside the droplet on receptors and components. After the self-alignment, the bumps are bridged to the contact pads. Adding mild acids to the water droplet

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prevents solder oxidation during capillary self-alignment.115 Other solvents that have been used for self-alignment include ethylene glycol.117Ionic liquids represent moreover an interesting and so far overlooked option. Ionic liquids are salts with low melting points (typically below 100 °C) and moderate electrical conductivity. They afford negligible volatility and a wide range of surface tension values depending on their chemical composition.118,119

Adhesives Adhesives have several attractive properties for capillary self-alignment.120 Many adhesives are liquid at room temperatures, and only after being heated up they cure to make a permanent bond. The downside is that many such adhesives have small surface tension, and are not easily confined inside receptors (Section 3.1). The problem can be solved by increasing the interfacial tension, for instance by immersing the adhesive droplet in an immiscible liquid. The interfacial tension of all hydrocarbons (such as hexadecane,27 besides adhesives) in contact with water increases,121 so that they can be confined within hydrophobic areas. Electrical connections in adhesive-based capillary self-alignment can also be established by using anisotropic conductive adhesives (ACAs),48,122 together with low-melting point alloys.122

Component Feeding Before capillary self-alignment can take place, the components need to be brought into contact with the liquid on the receptors. Three main strategies have been devised for component feeding: deterministic feeding, in series or in parallel, and parallel stochastic feeding.

For deterministic feeding, the components can be fed using a robotic pick-and-place device, either in series17,117,123 or in parallel.18,72,124 Parallel deterministic feeding includes laser-induced release and transfer from a donor substrate,125,126 roll-to-roll mechanisms,12,127 parallel component handling over temporary carriers92,124 and floatation.59,76 Feeding can be accomplished also stochastically, as often done in fluidic selfassembly.28,29,61,111,128,129 In this case, the components are introduced to a liquid bath and driven towards the receptors using fluidic flow and/or vibration to impose movement and overcome friction and component aggregation. The main feeding strategies are illustrated in Fig. 14 and their properties summarized in Table 3.

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Figure 14: Main component feeding techniques: (a) single-component and (b) parallel multi-component pick-and-place, using e.g. vacuum or capillary gripping; (c) laser-induced component release and transfer; (d) fluidic agitation and vibration feeding for fluidic self-assembly; (e) floatation feeding. Table 3: Comparison of component feeding techniques Method

Single component pick-andplace

Parallel pick-andplace

Laserinduced releasing

Fluidic feeding and vibration

Floatation feeding

Speed [unit per s]

High-speed: >15 lowspeed: 1000

Accuracy [μm]

High-speed: 4 unit per s) and a yield larger than 99% has been reported, without indicating the actual assembly accuracy. The intrinsic parallelism of fluidic self-assembly makes the process scalable to the integration of microcomponents onto arbitrarily large substrates, suitable

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for macroelectronic applications such as flexible substrates for solid-state lighting.

High precision Capillary self-alignment can easily achieve sub-micron registration accuracy.28,59–61,63,74 Whereas the highest absolute accuracy is normally achieved for sub-millimeter components, millimetre-sized components have also achieved similarly high relative accuracy. In ref. 113, droplets of diluted hydrofluoric acid have been used to align arrays of 5 × 5 mm2 chips onto hydrophilic receptors on a substrate (Fig. 16d). Average and best alignment accuracy were 400 nm and 50 nm, respectively. Dubey et al. have recently demonstrated accurate ( Cu de-mix. Only samples including Al particles < 20 µm remain mixed. These are the mixtures with both Al and Cu particles < 20 µm and those with Cu particles > 40 µm. The remaining fall into the intermediate category. During the experiments, it is observed that most mixtures categorized as de-mixed after vibration already show some degree of segregation after the mixing procedure. An exception is mixtures containing Cu particles 32–40 µm: they only de-mix during vibration. The mixtures ending up in the intermediate category appear fully mixed before vibration. These results confirm that creation of reliably stable Al-Cu powder mixtures is probably not possible with particle sizes > 20 µm that are established in LBM of Al without SiOx.

Figure 8: De-mixing of Al-Cu powder mixtures under variation of particle size fractions from test tube experiments, 94.7 wt% Al, 5 wt% Cu, and 0.3 wt% SiOx; (a) results; (b) example of powder remaining mixed; (c) example of intermediate powder with local Cu accumulations; (d) example of de-mixed powder with clearly visible segregation of Cu particles on top and bottom of the test tube.

To explain these observations, the dispersive coefficient of Cu, , is qualitatively compared between lower and higher Cu concentrations, qCu. is a measure for mobility of Cu particles among the majority of Al. Motion of particles is provided by vibration of the bulk powder. Decreasing with increasing qCu causes segregation [37]. Four cases are distinguished, as illustrated in Figure 9a–d.

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Figure 9: Considerations of changing mobility of Cu particles, D∗Cu, depending on local Cu concentration, qCu, and different particle sizes, in vibrated bulk; (a) Al and Cu < 20 µm; (b) Al > 20µm and Cu < 20 µm; (c) Al < 20 µm and Cu > 40 µm; (d) Al > 20 µm and Cu > 40 µm.

With all particles < 20 µm, the experimental result of no segregation remains constant with increasing qCu, as leads to the assumption that shown in Figure 9a. Al and Cu particles have similar size, but different mass. To explain the experimental result, it seems most plausible to assume that different mass has no significant effect in this case, which is characterized by small particle sizes < 20 µm and high flowability ensured by dry coating with SiOx nanoparticles. The second case of Al > 20 µm and Cu < 20 µm is illustrated in Figure 9b. The smaller Cu particles fit in gaps between the larger Al particles, which than among other Cu particles of similar size. Higher grants them higher qCu increases chances that near gaps between larger Al particles are occupied . Assuming mass has less by other Cu particles, which further decreases effects at this size of Cu particles is supported by the observation that Cu accumulations on top of test tubes are more pronounced than with larger Cu particles. Gravity and smaller particle size could be expected to result in downward transport of Cu under vibration [38,40,42].

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In the case of Al < 20 µm and Cu > 40 µm, which is illustrated in Figure 9c, the experimental results show no segregation, which implies no reduced at increased qCu. This seems plausible, considering that the smaller Al particles, which are much more numerous than Cu particles in the investigated mixtures with 95 wt% Al, should be able to move into gaps between larger Cu particles. This way, they can separate Cu particles from each other. Since mass is expected to have significant effects at this size of Cu particles, it is apparently compensated by the smaller size of Al particles. A similar observation of stability under motion without convective transport is described for a mixture of 2.5 mm large glass balls with 3.5 mm steel balls, at a mass density ratio of 2.6 [42]. The fourth considered case is that of intermediate de-mixing observed for Al particles > 20 µm and Cu particles of the same or larger size, as is slightly lower at higher illustrated in Figure 9d. Theory implies that qCu, but not sufficiently lower to cause strong de-mixing. Due to the smaller difference in particle size compared to Figure 9c, the separating effect is expected to be weaker. Since mass should be significant in these Cu particle sizes, it may be speculated that a higher probability for collisions between Cu particles would reduce their average mobility compared to lower qCu, which increases probability for Cu particles to collide with lighter Al particles of lower momentum. Results of test tube experiments with quaternary Al-Cu-Mg-Ti powder mixture are put together in Figure 10. The same color coding is used as with binary mixtures in Figure 8. In direct comparison, de-mixing is less pronounced in quaternary mixtures than in binary mixtures. Al particle size < 20 µm is the dominating factor for stable mixtures indicated by green squares. 23 out of 27 mixtures containing Al < 20 µm remain mixed. Only two of 53 mixtures without Al < 20 µm remain mixed. The reasons for the deviating results might be in the inaccuracies of the evaluation procedure. Powders are fractioned by sieving, which results in overlapping PSD, as further detailed in the Appendix A. After mixing and prior to vibration, no segregation of Ti6Al4V and AlMg50 can be found. After vibration, many mixtures show dark grey striations of Ti6Al4V or AlMg50 powder. Also, dark grey powder segregated on top or at the bottom of test tubes can be distinguished from light grey Al. Visual contrast is less than with reddish Cu powder, but clear enough for unambiguous categorization. Contrast could be improved by artificial illumination with light of wavelength ranges adapted to the specific reflectivity of the involved powders.

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10. De-mixing of Al-Cu-Mg-Ti powderunder mixtures, under variation of obtained particle size, Figure 10.Figure De-mixing of Al-Cu-Mg-Ti powder mixtures, variation of particle size, from obtained

from vibrated test tubes, constant powder92.4 composition 4 under wt% Cu, variation 3 wt% AlMg50, Figure 10: of Al-Cu-Mg-Ti powder mixtures, of0.14 vibrated testDe-mixing tubes, constant powder composition wt% Al, 492.4 wt%wt% Cu,Al, 3 wt% AlMg50, 0.14 wt% ; dominating factor forparticle stabilitysize is Al particle size < 20 µm. wt%0.3 Ti6Al4V, and 0.3 wt% SiOfactor Ti6Al4V, and wt% SiO ; dominating for stability is Al < 20 µm. x particle size, obtained from vibrated test tubes, constant powder composition 92.4 wt% Al, Microscopic 4 wt% Cu, wt% AlMg50, 0.14 wt% Ti6Al4V, and 0.3 wt% SiOx; 3.5. Light Microscopic Analysis of 3 Mixture Homogeneity in Thin Powder Layers 3.5. Light Analysis of Mixture Homogeneity in Thin Powder Layers dominating factor for stability is Al particle size < 20 µm. The results thin powder microscopy inside the LBM machine Realizer SLM 50 are refined Theofresults of thinlayer powder layer microscopy inside the LBM machine Realizer SLM 50 x

are in Figurerefined 11. As in a criterion to As compare the homogeneity of binary Al-Cu mixtures three different Figure 11. a criterion to compare the homogeneity of binarywith Al-Cu mixtures with three Mixtures with all particles sieved 20 µm without SiOx, as shown in Figure 6a. The mixture with larger Al shows highest Cu content at location e, which is the furthest from the turning point of the wiper. A second, less pronounced peak of Cu content of the mixture with larger Al is found at location e, closest to the turning point of the wiper. Based on the current results, it cannot be excluded that another combination of particle size and mass may compensate for the selectivity of convective transport. This would probably be a powder mixture that is not stable against de-mixing under vibration, as the mixture with Al < 20 µm and Cu 40–50 µm segregated. To better understand the details of de-mixing phenomena during spreading of thin layers, it might be revealing to capture the powder movements in front of the wiper, in videos resolving single particles. Layer thickness, wiper geometry, and wiper velocity appear as probable influencing factors.

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Micrographic Definition of LBM Process Maps Experimentally defined process windows for binary Al-Cu powder mixtures with constant composition of 94.7 wt% Al, 5 wt% Cu, and 0.3 wt% SiOx in three different PSD are visualized in Figure 12. Figure 12a–c show ρrel in colors shaded from red, for 90%, to dark green, for 100%. The dependence of Δxy is depicted along the horizontal and of v along the vertical axis. Other process parameters, P = 100 W, Δz = 30 µm, and dspot = 66 µm, are kept constant. The white lines in Figure 12a–c designate constant Evol, as defined in equation (6). Pink contours mark ρrel = 99%. For each PSD, 32 cube samples of 5 mm edge length have been built on filigree supports and analyzed micrographically for ρrel and details of defects. In Figure 12a–c, each cube is represented by a black dot. The colors in between dots are interpolated. Selected micrographs are shown in Figure 12d–h, each marked with a small black symbol (pentagon, triangle, circle or square), which tags the respective data point in Figure 12a–c. Three combinations of particle size fractions of Al and Cu are chosen, one with all particles < 20 µm, illustrated in Figure 12a,d,e, one with larger Cu, illustrated in Figure 12b,f,g, and one with larger Al particles, as shown in Figure 12c,h. With all three mixtures, ρrel > 99.5% have been achieved. Examples of such cubes are shown in Figure 12d,f,h. From the other micrographs, reasons for lower ρrel can be concluded. Figure 12e with ρrel = 88.6% shows voids elongated in parallel to Z of approximately constant width and distance. The used Δxy = 154 µm must be too large for reliable coalescence. Evol = 193 J/mm³ alone cannot give a satisfactory explanation, as lower 179 J/mm³ suffices for higher ρrel = 96.6%, at Δxy = 56 µm and v = 334 mm/s. An example for too high Evol = 1135 J/mm³ resulting in less than optimal ρrel = 98.8% is shown in Figure 12g. The spherical pores typically result from excessive energy input causing increased melt pool turbulence. As a potential alternative, they might result from entrapped hydrogen [80].

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Figure process maps of Al-Cu powderpowder mixturesmixtures with threewith different constant composition Figure12. 12.LBM LBM process maps of Al-Cu threePSD, different PSD, constant 94.7 wt% Al, 5 wt% Cu, and 0.3 wt% SiO ; (a–c) ρ over vx and ∆ ,relpink contour ρ xy = 99%, E isolines rel

x rel SiO ; (a–c)xy vol ρ over v and Δrel, pink contour ρ = different composition 94.7 wt% Al, 5 maps wt% Cu, and 0.3 wt% Figure 12: LBM of Al-Cu powder mixtures with three in white; Alprocess and particles 20 µm; (b) Al < 20 µm, Cu 32–40 (c) Al µm; isolines in Cu white; (a) Al