Extraction 2018 : proceedings of the first Global Conference on Extractive Metallurgy

This three volume set presents papers from the first collaborative global metallurgy conference focused exclusively on extractive topics, including business and economic issues. Contributions examine new developments in foundational extractive metallurgy topics and techniques, and present the latest research and insights on emerging technologies and issues that are shaping the global extractive metallurgy industry. The book is organized around the following main themes: hydrometallurgy, pyrometallurgy, sulfide flotation, and extractive metallurgy markets and economics.  Read more...

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Proceedings of the First Global Conference on Extractive Metallurgy

EDITORS

Boyd R. Davis • Michael S. Moats • Shijie Wang

CO-EDITORS Jochen Petersen Dean Gregurek Virginia S. T. Ciminelli Joël Kapusta Qian Xu Thomas P. Battle Ronald Molnar Mark E. Schlesinger Gerardo Raul Alvear Flores Jeff Adams Wenying Liu Evgueni Jak Niels Verbaan Graeme Goodall John Goode Michael L. Free Edouard Asselin Ian M. London Alexandre Chagnes Gisele Azimi David Dreisinger Alex Forstner Ronel Kappes Matthew Jeffrey Jaeheon Lee Tarun Bhambhani Graeme Miller

The Minerals, Metals & Materials Series

Boyd R. Davis ⋅ Michael S. Moats Shijie Wang ⋅ Dean Gregurek ⋅ Joël Kapusta Thomas P. Battle ⋅ Mark E. Schlesinger Gerardo Raul Alvear Flores ⋅ Evgueni Jak Graeme Goodall ⋅ Michael L. Free Edouard Asselin ⋅ Alexandre Chagnes David Dreisinger ⋅ Matthew Jeffrey Jaeheon Lee ⋅ Graeme Miller ⋅ Jochen Petersen Virginia S. T. Ciminelli ⋅ Qian Xu Ronald Molnar ⋅ Jeff Adams ⋅ Wenying Liu Niels Verbaan ⋅ John Goode ⋅ Ian M. London Gisele Azimi ⋅ Alex Forstner ⋅ Ronel Kappes Tarun Bhambhani Editors

Extraction 2018 Proceedings of the First Global Conference on Extractive Metallurgy

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Editors See next page

ISSN 2367-1181 ISSN 2367-1696 (electronic) The Minerals, Metals & Materials Series ISBN 978-3-319-95021-1 ISBN 978-3-319-95022-8 (eBook) https://doi.org/10.1007/978-3-319-95022-8 Library of Congress Control Number: 2018947474 © The Minerals, Metals & Materials Society 2018 This work is subject to copyright. All rights are reserved by the Publisher, whether the whole or part of the material is concerned, specifically the rights of translation, reprinting, reuse of illustrations, recitation, broadcasting, reproduction on microfilms or in any other physical way, and transmission or information storage and retrieval, electronic adaptation, computer software, or by similar or dissimilar methodology now known or hereafter developed. The use of general descriptive names, registered names, trademarks, service marks, etc. in this publication does not imply, even in the absence of a specific statement, that such names are exempt from the relevant protective laws and regulations and therefore free for general use. The publisher, the authors and the editors are safe to assume that the advice and information in this book are believed to be true and accurate at the date of publication. Neither the publisher nor the authors or the editors give a warranty, express or implied, with respect to the material contained herein or for any errors or omissions that may have been made. The publisher remains neutral with regard to jurisdictional claims in published maps and institutional affiliations. This Springer imprint is published by the registered company Springer Nature Switzerland AG The registered company address is: Gewerbestrasse 11, 6330 Cham, Switzerland

Boyd R. Davis Kingston Process Metallurgy Inc.

Jaeheon Lee University of Arizona

Michael S. Moats Missouri University of Science and Technology

Graeme Miller Miller Metallurgical Services Pty Ltd.

Shijie Wang Rio Tinto Kennecott Utah Copper Corporation

Jochen Petersen University of Cape Town

Dean Gregurek RHI Magnesita

Virginia S. T. Ciminelli Universidade Federal de Minas Gerais

Joël Kapusta BBA Inc.

Qian Xu Shanghai University

Thomas P. Battle Extractive Metallurgy Consultant

Ronald Molnar MetNetH20 Inc.

Mark E. Schlesinger Missouri University of Science and Technology

Jeff Adams Hatch

Gerardo Raul Alvear Flores Aurubis

Wenying Liu University of British Columbia

Evgueni Jak University of Queensland

Niels Verbaan SGS Minerals

Graeme Goodall XPS-Glencore

John Goode J.R. Goode and Associates Metallurgical Consulting

Michael L. Free University of Utah

Ian M. London Avalon Rare Metals Inc.

Edouard Asselin University of British Columbia

Gisele Azimi University of Toronto

Alexandre Chagnes University of Lorraine

Alex Forstner SGS Minerals

David Dreisinger University of British Columbia

Ronel Kappes Newmont Mining

Matthew Jeffrey Newmont Mining

Tarun Bhambhani Newmont Mining Corporation

Preface

While Extraction 2018 may seem like another new conference, it was actually born out of a membership initiative within three societies to reduce the number of extractive meetings in 2018. Set for August 26–29, 2018, in Ottawa, Canada, Extraction 2018 is the industry’s first collaborative global metallurgy conference focused exclusively on extractive topics. As such, we hope it will allow attendees to take stock of the current situation in extractive metallurgy, where issues around geography, resources, environment, and human capital are all in play and making for an ever-changing landscape. The conference has come to fruition out of the hard work from the staff of the three sponsoring societies. The Society for Mining, Metallurgy & Exploration (SME) has mobilized its team to manage the trade show and sponsorship, while The Minerals, Metals & Materials Society (TMS) has been instrumental in supplying the backbone of the conference for abstract management, author interaction, and marketing. The Metallurgy & Materials Society (MetSoc) of Canadian Institute of Mining, Metallurgy & Petroleum (CIM) organized the logistics for this conference while at the same time holding its first-ever annual meeting outside of Canada at Materials Science & Technology 2018 (MS&T18) in Columbus, Ohio. This opportunity has really allowed the extractive programming to come to Ottawa. The regular contact between societies has been beneficial for everyone and will have lasting positive impacts far after Extraction 2018 is over. A large number of volunteers from each society were also critical to the organization of the symposia. Each one of them has volunteered considerable personal time to help develop, review, and organize the conference. This includes short courses, tours, and posters, as well as the technical symposia. Extraction 2018 is the home for several important recurring symposia that examine new developments in foundational extractive metallurgy topics and techniques. It also offers new programming designed to share the latest research and insights on emerging technologies and issues that are shaping the global extractive metallurgy industry. The symposia include: 7th International Symposium on Advances in Sulfide Smelting; Peter Hayes Symposium on Pyrometallurgical Processing; Hydrometallurgy 2018;

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Gordon Ritcey Symposium: Advances in Hydrometallurgical Solution Purification Separations; Processing of Critical Metals; and finally, Sulfide Flotation. The response to the conference has been very encouraging—a reflection of the fact that there are actually a number of high-quality symposia being brought to the conference, but also perhaps an indication of the extractive community’s interest to meet once without the distraction of large shovels or nanomaterials. It was difficult to balance the number of abstract submissions with the desire to have a manageable conference, and hopefully, the right balance has been found. It will be a lot to take in, but we hope that everyone will get as much as they can from the conference. When the dust has settled, we will take stock of how things went, and if the stars realign in the future, we might try to do this again. But for now, welcome to Extraction 2018—you are among friends. Conference Co-chairs Boyd R. Davis, Kingston Process Metallurgy Inc. Michael S. Doats, Missouri University of Science and Technology Shijie Wang, Rio Tinto Kennecott Utah Copper Corporation

Contents

Part I

Pyrometallurgy Keynotes

Role of Research in Non-ferrous Metallurgy Development—Peter Hayes’ Contributions to Modern Pyrometallurgy . . . . . . . . . . . . . . . . . Phillip J. Mackey and Evgueni Jak The Role of Research in Pyrometallurgy Technology Development—From Fundamentals to Process Improvements—Future Opportunities . . . . . . . . . . . . . . . . . . . . . . . . . . Evgueni Jak Sulfide Smelting: Thirty-Five Years of Continuous Efforts to Find New Value Adding Solutions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . G. R. F. Alvear Flores, M. Löbbus, B. Springub, A. Fallah-Mehrjardi and A. Tappe The Changing World of Metallurgical Education . . . . . . . . . . . . . . . . . Peter C. Hayes Part II

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7th International Symposium on Advances in Sulfide Smelting

Sulfide Smelting Development in Japan During the Past Half Century . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Takahiko Okura and Hiromichi Takebe Review of Boliden Harjavalta Nickel Smelter . . . . . . . . . . . . . . . . . . . . Hannu Johto, Petri Latostenmaa, Esa Peuraniemi and Karri Osara Redesign and Rebuild of the Pan Pacific Copper Flash Smelting Furnace . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Glenn Stevens, Tatsuya Motomura, Tomoya Kawasaki, Misha Mazhar and Gary Walters

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Modelling Metallurgical Furnaces—Making the Most of Modern Research and Development Techniques . . . . . . . . . . . . . . . . . . . . . . . . . Evgueni Jak Pyrometallurgical Processing of Desulphurization Slags . . . . . . . . . . . . Christoph Pichler, Jürgen Antrekowitsch and Karl Pilz High Temperature Phase Formation at the Slag/Refractory Interphase at Ferronickel Production . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Christoph Sagadin, Stefan Luidold, Christoph Wagner and Alfred Spanring

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ISASMELT™ Technology for Sulfide Smelting . . . . . . . . . . . . . . . . . . . Ben Hogg, Stanko Nikolic, Paul Voigt and Paul Telford

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Refractory Design and the Role of Numerical Simulation . . . . . . . . . . . D. R. Kreuzer, C. Wagner, G. Unterreiter and J. Schmidl

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Wear Phenomena in Non-ferrous Metal Furnaces . . . . . . . . . . . . . . . . . D. Gregurek, C. Majcenovic, K. Budna, J. Schmidl and A. Spanring

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A Scientific Roadmap for Refractory Corrosion Testwork . . . . . . . . . . J. Schmidl, A. Spanring, D. Gregurek and K. Reinharter

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Investigation of Refractory Failure in a Nickel Smelting Furnace . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Wilson Pascheto, Roy Berryman, Robert Beaulieu and Maysam Moham

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Mathematical Modeling of Waterless Matte Granulator for Debottlenecking of Conventional Sulfide Smelters . . . . . . . . . . . . . . A. Navarra and F. Mucciardi

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Desulfurization of the Non-ferrous Smelter Flue Gases Based on Scrubbing with a Carbonate Eutectic Melt and Natural Gas Regeneration . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Valery Kaplan, Nurlan Dosmukhamedov and Igor Lubomirsky Advanced Thermochemical Fundamental and Applied Research to Improve the Integrity of the Steel Water Jacketed Furnace at Port Pirie . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . W. Watt, T. Hidayat, D. Shishin and E. Jak Sustainable Development Considerations in Primary Copper Smelting . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Krishna Parameswaran, Joe Wilhelm and Roberto Camorlinga Influence of Arsenic on the Chemical Wear of Magnesia-Chromite Refractories in Copper Smelting Furnaces . . . . . . . . . . . . . . . . . . . . . . . Katja Reinharter, Dean Gregurek, Christian Majcenovic, Jürgen Schmidl and Alfred Spanring

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Improved Copper Smelter and Converter Productivity Through the Use of a Novel High-Grade Feed . . . . . . . . . . . . . . . . . . . . . . . . . . . Eugene Jak, Denis Shishin, Will Hawker, James Vaughan and Peter C. Hayes Semi-discrete Dynamics and Simulation of Peirce-Smith Converting . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . A. Navarra, G. Lemoine, N. Zaroubi and T. Marin Development of Continuous Radar Level Measurement for Improved Furnace Feed Control . . . . . . . . . . . . . . . . . . . . . . . . . . . . Rodney Hundermark, Quintin van Rooyen, Paul van Manen, Chris Steyn, Afshin Sadri and David Chataway Research on Recovery of Valuable Metals in Waste Acid from Copper Smelting Flue Gas Acid-Making and Reduction and Harmless Treatment of Solid Wastes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Yan Wen, Zhen Bao and Xinmin Wu

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Fundamental Process Equilibria of Base and Trace Elements in the DON Smelting of Various Nickel Concentrates . . . . . . . . . . . . . . Pekka Taskinen, Katri Avarmaa, Hannu Johto and Petri Latostenmaa

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Challenges and Opportunities of a Lead Smelting Process for Complex Feed Mixture . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Christoph Zschiesche, Mehmet Ayhan and Jürgen Antrekowitsch

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Application of MPE Model to Nickel Smelting . . . . . . . . . . . . . . . . . . . Chunlin Chen Practice on Exploration of Oxygen-Enriched Converting Industrial Production by Kaldo Furnace . . . . . . . . . . . . . . . . . . . . . . . . Zhihua Wang

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Ust-Kamenogorsk Metallurgical Complex: A Silent Achiever . . . . . . . . Alistair Burrows and Turarbek Azekenov

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Trace Metal Distributions in Nickel Slag Cleaning . . . . . . . . . . . . . . . . Niko Hellstén, Pekka Taskinen, Hannu Johto and Ari Jokilaakso

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Case Study on the Application of Research to Operations—Calcium Ferrite Slags . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Stanko Nikolic, Denis Shishin, Peter C. Hayes and Evgueni Jak Kinetics of Oxidation of Pyrrhotite . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Anastasia Alksnis, Bo Li, Richard Elliott and Mansoor Barati Formation Mechanism of Ferronickel Alloy Due to the Reaction Between Iron and Nickeliferous Pyrrhotite at 850–900 °C . . . . . . . . . . Feng Liu, Mansoor Barati and Sam Marcuson

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Two-Step Copper Smelting Process at Dongying Fangyuan . . . . . . . . . Zhi-xiang Cui, Zhi Wang, Hai-bin Wang, Chuan-bing Wei, Peng Hou and Wu-zhao Du

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ISASMELT™—Flexibility in Furnace Design . . . . . . . . . . . . . . . . . . . . Stanko Nikolic, Ben Hogg and Paul Voigt

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Investigation of the Oxygen Bottom Blown Copper Smelting Process . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Qinmeng Wang and Xueyi Guo Top Submerged Lance Furnace Lining Cooling System Upgrade . . . . Allan MacRae and Brandon Steinborn Application Study on Technology of Reducing Copper Content in Discarded Slag . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Zhi-xiang Cui, Zhi Wang, Rui-min Bian, Chuan-bing Wei and Bao-jun Zhao Thermodynamic Consideration of Copper Matte Smelting Conditions with Respect to Minor Element Removal and Slag Valorization Options . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Eric Klaffenbach, Gerardo R. F. Alvear Flores, Muxing Guo and Bart Blanpain Optimizing Smelter Uptime Through Digital Asset Management . . . . . Bien Ferrer, Lucy Rodd, Adi Dhora, Richard MacRosty, Chris Walker and Mohamed Alhashme Reducing Refining Cycle Times to Extend Anode Furnace Campaign Life at Kennecott Copper . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Jun Enriquez, Ryan Walton, Adrian Deneys, Allen Chan, Bryan Bielec and Viktor Kilchyk Smelting Mechanism in the Reaction Shaft of a Commercial Copper Flash Furnace . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Zhou Jun and Chen Zhuo Progressing Towards Furnace Modernization by Utilizing Comparative Analysis of Acousto Ultrasonic-Echo (AU-E) Monitoring: Case Studies . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Afshin Sadri, Mitchell Henstock, Peter Szyplinski and Wai Lai Ying Kinetics of Roasting of a Sphalerite Concentrate . . . . . . . . . . . . . . . . . . Omid Marzoughi, Mohammad Halali, Davood Moradkhani and Christopher A. Pickles Thermodynamic Modeling of Oxygen Bottom-Blowing Continuous Converting Process . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Songsong Wang and Xueyi Guo

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Thermodynamic Considerations of Copper Complex Resources Smelting Process . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Miao Tian and Xueyi Guo Part III

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Peter Hayes Symposium on Pyrometallurgical Processing

Peter Charles Hayes’ Contributions to Metallurgical Research: Brief Biography and List of Publications . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Phillip J. Mackey and Evgueni Jak Integrated Pyrorefining of Lead at Teck’s Trail Operations . . . . . . . . . G. Richards and C. Curtis Relating Reported Carbon Dioxide Emissions to Iron and Steelmaking Process Details . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . P. Chris Pistorius Process Zones Observed in a 48 MVA Submerged Arc Furnace Producing Silicomanganese According to the Ore-Based Process . . . . . Joalet Dalene Steenkamp, Johan Petrus Gous, Wiebke Grote, Robert Cromarty and Helgard Johan Gous Heat Transfer to Copper Coolers in Freeze Lined Furnaces: The Role of Radiation and the Influence of Slag Liquidus . . . . . . . . . . M. W. Kennedy, A. MacRae and M. Shapiro Interfacial Reaction Between Magnesia Refractory and EAF Slag . . . . Jin Sung Han, Jung Ho Heo, Il Sohn and Joo Hyun Park

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Kinetics of Dephosphorization of Iron Carbon Alloys: The Importance of Competing Reactions, Slag Properties and CO Bubbles . . . . . . . . . . Kezhuan Gu, Phillip B. Drain, Brian J. Monaghan and Kenneth S. Coley

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A Study of Cementite Formation in the Reduction of Hematite by CO–CO2 Gas Mixture Using High Temperature XRD . . . . . . . . . . . Yury Kapelyushin, Yasushi Sasaki, Jianqiang Zhang and Oleg Ostrovski

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Condensation of SiO and CO in Silicon Production—A Literature Review . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Broggi Andrea and Tangstad Merete

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Phase Transformations from Quartz to Cristobalite . . . . . . . . . . . . . . . Karin Fjeldstad Jusnes, Merete Tangstad and Eli Ringdalen

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Towards Forming Micro-Bubbles in Liquid Steel . . . . . . . . . . . . . . . . . Roderick Guthrie and Mihaiela Isac

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Innovative Applications of Bubbles and Drops to Ferrous Process Technology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Youn-Bae Kang, Jungwook Cho, Sangjun Kim and Hae-Geon Lee

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Carbochlorination of Low-Grade Titanium Slag to Titanium Tetrachloride in Molten Salt . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Liang Li, Kaihua Li, Dachun Liu and Aixiang Chen Calciothermic Reduction and Electrolysis of Sulfides in CaCl2 Melt . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Ryosuke O. Suzuki, Nobuyoshi Suzuki, Yuta Yashima, Shungo Natsui and Tatsuya Kikuchi Reaction Behavior of Phosphorus in Multi-phase CaO–FeOX–SiO2–P2O5 Flux System . . . . . . . . . . . . . . . . . . . . . . . . . . . . Xiao Yang, Hiroyuki Matsuura and Fumitaka Tsukihashi Microanalysis and Experimental Techniques for the Determination of Multicomponent Phase Equilibria for Non-ferrous Smelting and Recycling Systems . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Taufiq Hidayat, Peter C. Hayes and Evgueni Jak Thermodynamic Study of the Equilibrium Distribution of Platinum Group Metals Between Slag and Molten Metals and Slag and Copper Matte . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Katsunori Yamaguchi High Temperature Recovery of Rare Earth Ortho-Ferrites from Permanent Magnets . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Muhamad Firdaus and M. Akbar Rhamdhani On the Evaporation of S from Liquid Fe–C–S Alloy . . . . . . . . . . . . . . . Youn-Bae Kang and Fahmi Tafwidli Extraction of Iron and Ferrosilicon Alloys from Low-Grade Bauxite Ores . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Jafar Safarian Kinetics of Bauxite Residue Sintering . . . . . . . . . . . . . . . . . . . . . . . . . . . Harrison Hodge, William Hawker, Peter Hayes and James Vaughan

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Multicomponent Thermodynamic Databases for Complex Non-ferrous Pyrometallurgical Processes . . . . . . . . . . . . . . . . . . . . . . . . Denis Shishin, Peter C. Hayes and Evgueni Jak

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Verification of Experimentally Determined Permeability and Form Coefficients of Al2O3 Ceramic Foam Filters (CFF) at High and Low Flow Velocity Using a CFD Model . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Massoud Hassanabadi, Mark W. Kennedy, Shahid Akhtar and Ragnhild E. Aune

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Computational Modeling of a Secondary Lead Reverberatory Furnace: Effect of Burden Geometry . . . . . . . . . . . . . . . . . . . . . . . . . . . Alexandra Anderson, Joseph Grogan, Gregory Bogin and Patrick Taylor

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Development of a Thermodynamic Database for the Multicomponent PbO-“Cu2O”-FeO-Fe2O3-ZnO-CaO-SiO2 System for Pyrometallurgical Smelting and Recycling . . . . . . . . . . . . . . . . . . . . . . . M. Shevchenko, P. C. Hayes and E. Jak Reduction in GHG Emission of Steel Production by Direct Injection of Renewable Biocarbon . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Ka Wing Ng, Louis Giroux and Ted Todoschuk Preparation of Ferronickel from Nickel Laterite Ore via Semi-molten Reduction Followed by Magnetic Separation . . . . . . . . . . . . . . . . . . . . . Xueming Lv, Lunwei Wang, Zhixiong You, Jie Dang, Xuewei Lv, Guibao Qiu and Chenguang Bai

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Thermodynamic Modeling of the Solid State Carbothermic Reduction of Chromite Ore . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Omid Marzoughi and Christopher A. Pickles

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Production of Ferromanganese Alloys from Silicomanganese Sludge and an Iron Source . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . M. Wallin, K. E. Ekstrøm and G. Tranell

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New Paradigms for Iron Ore Pelletization . . . . . . . . . . . . . . . . . . . . . . . S. K. Kawatra In Situ Micro Raman Study of the NO3− Electrochemical Behavior in Molten NaNO3–KNO3 Mixtures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Zhuo Sheng, Liang Li and Kaihua Li Dissolution of Graphite in Iron Manganese Alloys . . . . . . . . . . . . . . . . H. Kaffash and M. Tangstad Experimental Investigation of Pyrometallurgical Treatment of Zinc Residue . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Minna Rämä, Ari Jokilaakso, Lassi Klemettinen, Justin Salminen and Pekka Taskinen Dynamic Modelling of Molten Slag-Matte Interactions in an Industrial Flash Smelting Furnace Settler . . . . . . . . . . . . . . . . . . Nadir Ali Khan and Ari Jokilaakso

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High Temperature Characteristics of Slags Originating from the Production of Synthetic Tantalum Concentrate . . . . . . . . . . . 1007 Dominik Hofer, Stefan Luidold, Frank Schulenburg and Tobias Beckmann Recovery of Nickel and Vanadium from Heavy Oil Residues Using DC Plasma Smelting . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1017 Tim P. Johnson

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Iron Segregation Roasting Processes . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1029 P. Kerr, Q. Liu and T. H. Etsell Towards a Microwave Metal Extraction Process . . . . . . . . . . . . . . . . . . 1039 C. A. Pickles and O. Marzoughi The Influence of Aluminum on Indium and Tin Behaviour During Secondary Copper Smelting . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1061 Katri Avarmaa and Pekka Taskinen Behavior of Nickel as a Trace Element and Time-Dependent Formation of Spinels in WEEE Smelting . . . . . . . . . . . . . . . . . . . . . . . . 1073 Lassi Klemettinen, Katri Avarmaa, Pekka Taskinen and Ari Jokilaakso The Distribution of Sn Between CaO–CuOx–FeOy–SiO2 Slag and Copper Metal at 1300 °C . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1083 A. Van den Bulck, S. Turner, M. Guo, A. Malfliet and B. Blanpain Gaseous Reduction of Mn Ores in CO-CO2 Atmosphere . . . . . . . . . . . 1093 T. A. Larssen, M. Tangstad and I. T. Kero Optimization of Slag Composition in View of Iron Recovery and Dephosphorization in EAF Process . . . . . . . . . . . . . . . . . . . . . . . . . 1103 Jung Ho Heo and Joo Hyun Park A New Pyrometallurgical Recycling Technique for Lead Battery Paste Without SO2 Generation—A Thermodynamic and Experimental Investigation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1109 Yun Li, Yongming Chen, Chaobo Tang, Shenghai Yang, Lassi Klemettinen, Minna Rämä, Xingbang Wan and Ari Jokilaakso Understanding Viscosity-Structure Relationship of Slags and Its Influence on Metallurgical Processes . . . . . . . . . . . . . . . . . . . . . . . . . . . 1121 Tae Sung Kim, Jung Ho Heo, Jin Gyu Kang, Jin Sung Han and Joo Hyun Park Reduction of Manganese Ore Pellets in a Methane-Containing Atmosphere . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1129 Richard Elliott and Mansoor Barati Kinetics of Reduction-Carburization of Synthetic (Fe,Mg)(Cr,Al)2O4 Solid Solutions by Ar–CH4–H2 Gas Mixtures . . . . . . . . . . . . . . . . . . . . 1141 Vincent Canaguier and Leiv Kolbeinsen Preliminary Experimental Study of the Thermal Stability and Chemical Reactivity of the Phosphate-Based Binder Used in Al2O3-Based Ceramic Foam Filters (CFFs) . . . . . . . . . . . . . . . . . . . . . . 1153 Cathrine K. W. Solem, Robert Fritzsch and Ragnhild E. Aune

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Part IV

xvii

Hydrometallurgy Keynotes

The Evolution of Cobalt–Nickel Separation and Purification Technologies: Fifty Years of Solvent Extraction and Ion Exchange . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1167 Kathryn C. Sole Minimizing the Hydro in Hydrometallurgy . . . . . . . . . . . . . . . . . . . . . . 1193 G. T. Lapidus Part V

Hydrometallurgy 2018

Hydrometallurgical Extraction of Lead in Brine Solution from a TSL Processed Zinc Plant Residue . . . . . . . . . . . . . . . . . . . . . . . 1205 Rajiv R. Srivastava, Jae-chun Lee, Tam Thi Nguyen, Min-seuk Kim and Jingu Kang Alkali Metals Removal from Radioactive Wastewater by Combined CO2 Capture and Adsorption onto Bone Char . . . . . . . . . . . . . . . . . . . 1213 Elbert M. Nigri, André L. A. Santos, Leonardo F. Santos and Sônia D. F. Rocha Synthesis and Properties Characterization of Crystalline Polyferric Sulfate Adsorbent Used for Treating High As(III)-content Contaminated Water . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1225 Pingchao Ke and Zhihong Liu Custom Fiberglass Reinforced Plastic Piping (FRP) Applications in Mineral Processing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1233 David J. Chapman and Anthony M. Zacharewych Electrochemical Behavior of Chalcopyrite in Presence of Sodium Peroxodisulfate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1245 Hojat Naderi and Jochen Petersen Alternative Lixiviant for Copper Leaching from Chalcopyrite Concentrate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1257 Junmo Ahn, Jiajia Wu and Jaeheon Lee Hydrometallurgical Processing of Copper-Arsenic Concentrates . . . . . 1267 Jan Smit, Kelvin Buban, Mike Collins and Preston Holloway Sustainable Development Considerations in Copper Hydrometallurgy . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1279 Krishna Parameswaran, Chris Mapes, Aaron Ibarra, Justin Landrum and Tracy Morris Improved Process for Leaching Refractory Copper Sulfides with Hydrogen Peroxide in Aqueous Ethylene Glycol Solutions . . . . . . . . . . 1289 Ángel Ruiz-Sánchez and Gretchen T. Lapidus

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Copper Recovery from the Mine Tailings by Combination of Flotation with High-Pressure Oxidative Leaching and Solvent Extraction . . . . . . 1299 Atsushi Shibayama, Baisui Han, Kazutoshi Haga, Zoran Stevanović, Radojka Jonović, Ljiljana Avramović, Radmila Marković, Daniela Urosević, Yasushi Takasaki, Nobuyuki Masuda and Daizo Ishiyama Douglas Centenary Commemoration 1918–2018: Engineering the Science—James Douglas, Early Hydrometallurgy and Chile . . . . . 1309 William W. Culver Making the Right Selection: A Comparative Analysis for the Treatment of Refractory Gold Concentrates . . . . . . . . . . . . . . . . . . . . . 1327 Rodney Clary, Paul DiNuzzo, Thomas Hunter and Saleem Varghese Evolution of Metallurgical Parameters at Mantoverde Heap Leach Operation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1339 Gabriel Zarate The Effect of Aeration on Chalcocite Heap Leaching . . . . . . . . . . . . . . 1353 Wenying Liu and Giuseppe Granata Study of the Diffusion of Cu(II) as an Oxidant Through Simulated Particle Pores in a Novel Model Apparatus . . . . . . . . . . . . . . . . . . . . . . 1361 B. Manana, J. Petersen and R. Ram Filtration Properties of Ferric Hydroxide Precipitate in Nickel Production . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1373 Ina Beate Jenssen, Mona Aufles Hines, Ole Morten Dotterud, Oluf Bøckman and Jens-Petter Andreassen The Effects of Experimental Variables on Iron Removal from Nitrided Malaysian Ilmenite by Becher Process . . . . . . . . . . . . . . . . . . . 1383 Eltefat Ahmadi, Noor Izah Shoparwe, Najwa Ibrahim, Sheikh Abdul Rezan Sheikh Abdul Hamid, Norlia Baharun, Kamar Shah Ariffin, Hashim Hussin and M. N. Ahmad Fauzi Water: An Increasingly Valuable and Challenging Resource for the Mining and Metallurgical Industry to Manage Effectively . . . . 1397 V. Ram Ramachandran Development of an Encapsulation Process to Extend the Stability of Scorodite Under Wider pH and Redox Potential Range Conditions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1411 Fuqiang Guo and George P. Demopoulos Arsenic Removal from Arsenic Bearing Materials Produced from Metallurgical Processes of Copper, Lead and Zinc . . . . . . . . . . . . 1421 Zhihong Liu, Zhiyong Liu, Jianxin Zhang, Siwei Li and Yuhu Li

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Modular Reactors and Utilization in Small Scale Direct Leaching Zinc Plant Expansions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1433 Björn Saxen, Tuomas Hirsi, Teemu Ritasalo and Marko Lahtinen Jarogain Process—A Hydrometallurgical Option to Recover Metal Values from RLE Zinc Residue and Steel Dust . . . . . . . . . . . . . . . . . . . 1443 Pertti Koukkari, Petteri Kangas, Mari Lundström, Sami Kinnunen, Jussi Rastas and Pekka Saikkonen The Betts Process at Trail Operations—History of Invention and Experiences from Over a Century of Operation . . . . . . . . . . . . . . . 1455 S. Fitzel and J. A. Gonzalez Building a Cloud-Based Operator Training Simulation Software for Pressure Oxidation Process . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1465 Mikko Loponen and Kristian Lillkung An Approach to Evaluate the Effect of Organic Compounds (Impurities and Additives) on Metal Electrowinning . . . . . . . . . . . . . . . 1473 D. Majuste, V. S. T. Ciminelli, P. R. Cetlin, E. L. C. Martins and A. D. Souza A Comprehensive Model for Metal Electrowinning Processes . . . . . . . . 1485 Zongliang Zhang, Joshua M. Werner and Michael L. Free Characterizing the Role of Organic Additives in Copper Electrowinning . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1497 C. Coetzee, M. Tadie and C. Dorfling Study of Electrochemical Behaviour and Surface Morphology of Copper Electrodeposit from Electrorefining with Lignin-Based Biopolymer and Thiourea as Additives . . . . . . . . . . . . . . . . . . . . . . . . . . 1509 M. Z. Mubarok, R. A. Lauten, R. Ellis, D. Ramdani and M. Syaifudin Manganese-Chloride Interactions on Pb–Ag Anode Behaviour in Synthetic Sulfuric Acid Electrolytes . . . . . . . . . . . . . . . . . . . . . . . . . . 1521 Charles E. Abbey, Wei Jin and Michael S. Moats Pb–Ca–Sn Anode Potential as a Function of Cobalt, Iron and Manganese in Synthetic Sulfuric Acid Electrolytes . . . . . . . . . . . . . 1535 C. E. Abbey and M. S. Moats The Effect of Anodic Potential on Surface Layers of Chalcopyrite during Ammonia–Ammonium Chloride Leaching . . . . . . . . . . . . . . . . . 1547 Xiaoming Huaa, Yongfei Zheng, Qian Xu, Xionggang Lu, Hongwei Cheng, Xingli Zou, Qiushi Song, Zhiqiang Ning and Michael L. Free

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Treatment of Arsenic Sulfide Sludge for Arsenic Stabilization and Copper Extraction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1555 Jianxin Zhang and Zhihong Liu Study of the Rate Controlling Steps in the Removal of Magnesium Impurities in Hydrogen Assisted Magnesiothermic Reduction of TiO2 by Leaching . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1567 Sayan Sarkar, Wyatt McNeill, Jayson Benedict and Michael L. Free EMD Deposition on Mn2O3/Ti Anode for Manganese Recovery from Zinc Electrowinning Solutions . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1575 Yuuki Yoshida, Kenji Kawaguchi and Masatsugu Morimitsu Pilot Plant Commissioning and Operations for Copper Sulfide In Situ Indirect Bioleaching . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1583 Theodore Ineich, Jan Kwak, Mark Damhuis, Bart Zaalberg, Doris Hiam-Galvez and Wickus Slabbert The Dewatering Behaviour of Transformed Ferri-Oxyhydroxide Precipitates Formed Under Moderate Temperature and Varying Fe(III) Concentrations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1597 Cledwyn Mangunda, Jochen Petersen and Alison Lewis The Hydrothermal Reaction and Kinetic of Enargite . . . . . . . . . . . . . . 1611 Gerardo Fuentes In Situ Precipitation of Scorodite in Atmospheric Leaching of Enargite . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1621 Fazel G. Jahromi and Ahmad Ghahreman Solubility of Rare Earth Salts in Sulphate-Phosphate Solutions of Hydrometallurgical Relevance . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1631 A. M. T. S. Bandara, G. Senanayake, D. I. Perera and S. Jayasekera Pressure Oxidation of Enargite Concentrates Containing Gold and Silver . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1645 Kimberly Conner-Mills and Corby Anderson Emulsion Mediated Low Temperature Pressure Leaching of Base Metals from Mixed Sulfide Minerals Through Enhanced Oxygen Mass Transfer . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1661 Shivendra Sinha, Devabrata Mishra, Saurabh Shekhar, Archana Agrawal and Kamla Kanta Sahu Equipment Selection for Chloride Circuits . . . . . . . . . . . . . . . . . . . . . . . 1671 G. R. Waters and G. B. Harris Understanding Cyanidation of Silver from Batch and Continuous Medium Temperature Pressure Oxidation Generated Residues . . . . . . 1685 R. T. Seaman and K. Mayhew

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Effect of DO, Free Cyanide and Mineralogy on Gold Cyanidation Mechanism: An Electrochemical and Surface Analysis Study . . . . . . . . 1697 Rina Kim, Ahmad Ghahreman and Michel Epiney Activation and Deactivation Effects of Lead on Gold Cyanidation . . . . 1709 Rina Kim, Ahmad Ghahreman and Michel Epiney Australian Hydrometallurgy Research and Development . . . . . . . . . . . 1721 James Vaughan, Weng Fu, Hong Peng, Will Hawker, Peter C. Hayes and Dave Robinson Dezincing of Galvanized Steel by Sulfuric Acid Leaching . . . . . . . . . . . 1733 J. Grogan, G. M. Martins and C. G. Anderson Extraction of the Surface-Coated Metals from Waste Acrylonitrile Butadiene Styrene Plastics in an Ammoniacal Solution . . . . . . . . . . . . . 1743 Minji Jun, Tae Gyun Kim, Jae-chun Lee, Rajiv R. Srivastava and Min-seuk Kim Comparative Evaluation of Sulfuric and Hydrochloric Acid Atmospheric Leaching for the Treatment of Greek Low Grade Nickel Laterites . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1753 Christiana Mystrioti, Nymphodora Papassiopi, Anthimos Xenidis and Konstantinos Komnitsas Novel Process for Comprehensive Utilization of Iron Concentrate Recovered from Zinc Kiln Slag . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1765 Zhi-yong Liu, Huan Ma, Zhi-hong Liu and Qi-hou Li Selection of Microorganism for the Bio-Oxidation of a Refractory Gold-Concentrate with Focus on the Behaviour of Antimony Sulphides . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1777 Liliane C. Carvalho, Suzimara R. Silva, Romeu M. N. Giardini, Lucas S. Magalhães, Michael L. M. Rodrigues and Versiane A. Leão Gold Leaching by Sodium Chloride and Calcium Hypochlorite Solutions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1787 Felipe A. de Carvalho, Andrea Resende and Versiane A. Leão The Kinetic of Atmospheric Acid Leaching of Brazilian Lateritic Nickel . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1797 André L. A. Santos, Emily M. A. Becheleni, Rísia M. Papini, Paulo R. M. Viana and Sônia D. F. Rocha Hydrolytic Precipitation of Nanosized TiO2 Phases for Use as Photocatalytic Sorption Media in Effluent Treatment . . . . . . . . . . . . . . 1809 Konstantina Chalastara, Fuqiang Guo and George P. Demopoulos

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Oxygen Pressure Leaching Behavior of Nickel from Black Shale . . . . . 1819 Zhigan Deng, Xingbin Li, Minting Li, Chang Wei, Gang Fan and Cunxiong Li Part VI

Gordon Ritcey Symposium: Advances in Hydrometallurgical Solution Purification Separations

The Feasibility of Separation of Rare Earth Elements by Use of Electrodialysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1831 Sanaz Mosadeghsedghi, Saviz Mortazavi and Maziar E. Sauber Separation of Cobalt and Metals in Acidic Chloride Solutions Using Diffusion Dialysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1839 Zizheng Zhou and David Dreisinger Application of Donnan Dialysis to the Separation and Recovery of Cations During Hydrometallurgical Recycling of Lithium Ion Batteries . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1855 Alexandru C. Sonoc and Jack Jeswiet Electrolytic Salt Splitting for Sulfuric Acid and Caustic Recovery: Can It Be Cost-Effective? . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1867 Alexander Burns and Clive Brereton The Industrial Application of Ultrafiltration and Reverse Osmosis for the Recovery of Copper, Silver and Cyanide from Gold Leach Liquors . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1883 Farhang Hedjazi and A. John Monhemius Extraction of Water from Contaminated Effluents by Forward Osmosis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1893 Georgios Kolliopoulos and Vladimiros G. Papangelakis Preliminary Test Results on High Rate Compressible Media Filtration Promise Benefits for Leach Liquor Clarification . . . . . . . . . . 1903 A. Galvan A Review on Application of Crown Ethers in Separation of Rare Earths and Precious Metals . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1913 V. I. Lakshmanan and S. Vijayan Cadmium and Nickel Adsorption Study Using Modified Biosorbent . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1931 Peijia Lin and Jaeheon Lee Imprinted Resin—The 21st Century Adsorbent . . . . . . . . . . . . . . . . . . . 1943 Sue Ritz, Jon Gluckman, Glen Southard, Brandi Maull and Dae Jung Kim

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Nickel Recovery from Hyperaccumulator Plants Using a Chelating Resin . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1961 Mathilde Guilpain, Baptiste Laubie and Marie-Odile Simonnot Advances in the Development of Electrostatic Solvent Extraction for Process Metallurgy . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1971 Don Ibana, Simon Assmann and Marc Steffens Experimental Study and CFD Simulation of a Solvent Extraction Pulsed Column with Novel Ceramic Internals . . . . . . . . . . . . . . . . . . . . 1979 Heng Yi, Weiyang Fei and Geoffrey W. Stevens ®

Outotec’s VSF X Implementation to Provide Modular Solvent Extraction Technology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1987 Tuomas Hirsi and Rami Saario Settling Behavior and CFD Simulation of a Gravity Separator . . . . . . 1997 Jan Steinhoff and Hans-Jörg Bart Equilibrium Modeling of Solvent Extraction and Stripping of Copper (II), Nickel(II), and Ammonia for Ammoniacal Process ® Using LIX 84-I . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2009 Shubin Wang, Jie Li, Hirokazu Narita and Mikiya Tanaka Modelling Synergistic Solvent Extraction of Nickel and Cobalt . . . . . . 2017 Mike Dry Liquid-Liquid Extraction of Cobalt(II), Nickel(II) and Manganese(II) from Acidic Chloride Media . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2027 Alexandre Chagnes and Kateryna Omelchuk Technological Advances, Challenges and Opportunities in Solvent Extraction from Energy Storage Applications . . . . . . . . . . . . . . . . . . . . 2033 Laurent Cohen, Tyler McCallum, Owen Tinkler and William Szolga Separation and Purification of Value Metals from Aqueous Chloride Solutions by Solvent Extraction . . . . . . . . . . . . . . . . . . . . . . . . 2047 V. I. Lakshmanan, R. Sridhar, D. Tait, R. deLaat, M. A. Halim and J. Chen Recovery of Palladium from Spent Catalysts—A Critical State-of-the-Art Review . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2063 Ana Paula Paiva Unique Hydrometallurgical Process for Copper-Anode Slime Treatment at Saganoseki Smelter and Refinery . . . . . . . . . . . . . . . . . . . 2075 Takahiro Furuzono, Atsushi Fujimoto, Tomohisa Takeuchi and Kazuaki Takebayashi

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Contents

Separation of Lead from Chalcopyrite Slurry Using Resin-in-Pulp . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2085 Weng Fu, Sabrina Lao, Yepeng Ding and James Vaughan Ion Exchange Resin—Pilot and Resin Testing . . . . . . . . . . . . . . . . . . . . 2093 Donald D. Downey Antimony and Bismuth Control in Copper Electrolyte by Ion Exchange . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2107 Katerina Kryst and Phillip (Rocky) Simmons Development and Screening of Resins to Recover REE and Scandium from Different Sources . . . . . . . . . . . . . . . . . . . . . . . . . . 2113 Mikhail Mikhaylenko Green Chemistry Principles Applied to the Selective Separation and Purification of Specialty Metals Using Molecular Recognition Technology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2123 Steven R. Izatt, Ronald L. Bruening, Neil E. Izatt and Reed M. Izatt Breakthrough in Uranium Recovery from Saline Liquors by Ion Exchange . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2137 Karin Soldenhoff, James E. Quinn, Tomasz Safinski, Keith Bowes and Merrill Ford The Use of Ion Exchange to Improve Revenue via the Removal of Impurities . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2149 Johanna van Deventer and Yoshinari Mori Selective Separation of Iron from Simulated Nickel Leach Solutions Using Ion Exchange Technology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2161 René A. Silva, Yahui Zhang, Kelly Hawboldt, Lesley James and Wesley Saunders Recovery of Copper Nanoparticles from AMD by Cementation with Iron and SDS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2173 G. Granata and C. Tokoro Separation of Uranium and Molybdenum in U–Mo Ore Alkaline Leach Liquor . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2183 Meifeng Zhi, Zujun Shu, Fengqi Zhao, Zhiquan Zhou, Yongming Zhang and Xiaohao Cao Separation of Iron (III) and Nickel (II) from Acidic Sulfate Leaching Solution of Molybdenum-Nickel Black Shale . . . . . . . . . . . . . 2193 Zhigan Deng, Xingbin Li, Gang Fan, Chang Wei, Cunxiong Li and Minting Li

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Part VII

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Processing of Critical Metals

Material Criticality: Comparing China, the EU, Japan and the USA . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2205 Roderick G. Eggert Lithium Extraction and Utilization: A Historical Perspective . . . . . . . . 2209 I. Peerawattuk and E. R. Bobicki Extraction of Lithium from Brine—Old and New Chemistry . . . . . . . . 2225 Mike Dry ®

Development of SiLeach Technology for the Extraction of Lithium Silicate Minerals . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2235 C. S. Griffith, A. C. Griffin, A. Roper and A. Skalski Lithium and Boron Extraction from the Rhyolite Ridge Ore Nevada USA . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2247 Peter Ehren Assessment of Lithium Pegmatite Ore Bodies to Determine Their Amenability to Processing for the Extraction of Lithium . . . . . . . . . . . 2261 Mark G. Aylmore Revisiting the Traditional Process of Spodumene Conversion and Impact on Lithium Extraction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2281 Colin Dessemond, Francis Lajoie-Leroux, Gervais Soucy, Nicolas Laroche and Jean-François Magnan Flowsheet Development for Benefication of Lithium Minerals from Hard Rock Deposits . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2293 M. Oliazadeh, M. Aghamirian, S. Ali, E. Legault and C. Gibson Hydrometallurgical Extraction of Rare Earth Elements from Coal . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2309 Rick Honaker, Xinbo Yang, Alind Chandra, Wencai Zhang and Joshua Werner Rapid and Selective Leaching of Actinides and Rare Earth Elements from Rare Earth-Bearing Minerals and Ores . . . . . . . . . . . . 2323 Laurence Whitty-Léveillé, Nicolas Reynier and Dominic Larivière Supercritical Fluid Extraction of Rare Earth Elements from NiMH Battery . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2329 Y. Yao, J. Zhang, J. Anawati and G. Azimi Extraction and Purification of Rare Earth Elements and Cobalt from NdFeB Magnet Wastes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2337 Hoda Emami and Pouya Hajiani

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Sulfuric Acid Baking and Leaching of Rare Earth Elements, Thorium and Phosphate from a Monazite Concentrate . . . . . . . . . . . . . . . . . . . . 2343 John Demol, Elizabeth Ho and Gamini Senanayake Development of a Metallurgical Process for Eramet’s Mabounié Nb-REE Project . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2353 G. Nazari, J. Lamotte, M. Ries, J. Agin, E. Tizon, S. Kashani-Nejad, M. Bellino and B. Krysa SCRREEN: Solutions for Critical Raw Materials—A European Expert Network . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2367 Stéphane Bourg and co-authors from the SCRREEN Project’s Partners Selective Extraction of Rare Earth Elements from Complex Monazite Ores . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2381 Leandro Augusto Viana Teixeira, Ruberlan Gomes Silva, Daniel Majuste and Virginia Ciminelli Leaching Kinetics of Rare-Earth Elements from Complex Ores by Acidic Solutions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2391 Hüseyin Eren Obuz, Haydar Günes, Aslıhan Kara, Dilan Ugurluer, Yurdaer Babuccuoglu and Murat Alkan Selective Extraction and Recovery of Rare Earth Metals (REMs) from NdFeB Magnet Grinding Sludge . . . . . . . . . . . . . . . . . . . . . . . . . . 2399 Waraporn Piyawit, Pisit Sawananusorn, Loeslakkhana Srikhang, Panya Buahombura, Narong Akkarapattanagoon, Tapanee Patcharawit and Sakhob Khumkoa Recovery of Phosphorous and Rare Earth Elements from an Apatite Concentrate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2409 Mahmood Alemrajabi, Kerstin Forsberg and Åke Rasmuson Hydrothermal Modification of Phosphogypsum to Improve Subsequent Recovery of Rare Earths . . . . . . . . . . . . . . . . . . . . . . . . . . . 2415 V. Yahorava, E. Lakay, W. Clark and J. Strauss Study of the Deportment of REEs in Ion Adsorption Clays Towards the Development of an In Situ Leaching Strategy . . . . . . . . . 2429 Cody Burcher-Jones, Sfiso Mkhize, Megan Becker, Rahul Ram and Jochen Petersen Thermodynamics of Tungsten Ores Decomposition Process Options . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2441 Leiting Shen, Xiaobin Li and Pekka Taskinen Recovery of Tungsten from Spent V2O5–WO3/TiO2 Catalyst . . . . . . . . 2455 In-Hyeok Choi, Gyeonghye Moon, Jin-Young Lee and Rajesh Kumar Jyothi

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Scrap Recycling of Tungsten-Based Secondary Material for the Recovery of Tungsten Monocarbide (WC) and Other Valuable Constituents Using an Acid Leach Process: A Preliminary Study . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2471 A. Shemi, S. Ndlovu and N. Sacks Solubility of Germanium Dioxide in Commonly Used Acids—Effect of Acid Strength, Temperature, and Water Activity . . . . . . . . . . . . . . . 2481 T. Feldmann, S. Nosrati and F. Bélanger Research on the Behavior of Germanium in the Leaching Process of Germanium-Bearing Zinc Oxide by Sulfuric Acid . . . . . . . . . . . . . . 2493 Tao Jiang and Zhihong Liu Natural Graphite Purification Through Chlorination in Fluidized Bed Reactor . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2505 K. Adham and G. Bowes Critical Materials Traceability: More Important Than Metallurgy . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2513 J. R. Goode A Process Flowsheet for the Extraction of Niobium, Titanium, and Scandium from Niocorp’s Elk Creek Deposit . . . . . . . . . . . . . . . . . 2523 Niels Verbaan, Mike Johnson, Tassos Grammatikopoulos, Eric Larochelle, Scott Honan, Kelton Smith and Rick Sixberry Recycling of Li-Ion and Li-Solid State Batteries: The Role of Hydrometallurgy . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2541 François Larouche, George P. Demopoulos, Kamyab Amouzegar, Patrick Bouchard and Karim Zaghib Hydrothermal Production of Lithium Metal Silicate Powders with Controlled Properties for Application to Li-ion Batteries . . . . . . . 2555 Yan Zeng, Karim Zaghib and George P. Demopoulos Production of Purified Lithium Salts with a One-Stage Solid Phase Extraction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2565 R. Hammen Electrochemical and Diffusion Assisted Dispersion Methods for Lithium-7 Enrichment from Liquid Media . . . . . . . . . . . . . . . . . . . 2575 Prashant K. Sarswat and Michael L. Free Fundamental Understanding of the Flotation Chemistry of Rare Earth Minerals . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2581 Stephanie Trant, Greer Galt, Avimanyu Das and Courtney A. Young

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Simulation of the Flotation of Bear Lodge Ore and a Preliminary Economic Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2597 H. Cui and C. Anderson Efficient Recovery of Neodymium from Neodymium–Iron–Boron Magnet . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2609 Jiakai Zhang, Feixiong Zhang and Gisele Azimi Recovery of REEs from End-of-Life Permanent Magnet Scrap Generated in WEEE Recycling Plants . . . . . . . . . . . . . . . . . . . . . . . . . . 2619 Sebastiaan Peelman, Prakash Venkatesan, Shoshan Abrahami and Yongxiang Yang A Comparison Between Batch and Continuous Processes in Impurity Removal from REE Water Leach Solution by Lime and Limestone Neutralization . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2633 Farzaneh Sadri, Zhi Yang and Ahmad Ghahreman Innovative Coupled Hydrometallurgical and Pyrochemical Processes for Rare Earth Recycling . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2647 V. Blet, E. Andreiadis, J. Serp and M. Miguirditchian REE Recovery from the Fern D. Dichotoma by Acid Oxalic Precipitation After Direct Leaching with EDTA . . . . . . . . . . . . . . . . . . 2659 Baptiste Laubie, Zeinab Chour, Ye-Tao Tang, Jean-Louis Morel, Marie-Odile Simonnot and Laurence Muhr Extraction of Cobalt and Nickel from a Pyrrhotite Rich Tailings Sample via Bioleaching . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2669 R. Cameron, B. Yu, C. Baxter, A. Plugatyr, R. Lastra, M. Dal-Cin, P. H. J. Mercier and N. Perreault Vanadium Extraction from Low Concentrated Iron Bearing Sources by a New Method . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2681 B. Nowak, R. Stastny and H. Weissenbaeck Recovery of Lithium from the Great Salt Lake Brine . . . . . . . . . . . . . . 2695 Rajashekhar Marthi and York R. Smith Crystallization of a Pure Scandium Phase from Solvent Extraction Strip Liquors . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2707 Edward Peters, Carsten Dittrich, Serif Kaya and Kerstin Forsberg Innovative and Sustainable Valorization Process to Recover Scandium and Rare Earth Elements from Canadian Bauxite Residues . . . . . . . . . 2715 John Anawati, Sable Reid and Gisele Azimi

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Extraction of Scandium (Sc) Using a Task-Specific Ionic Liquid Protonated Betaine Bis(Trifluoromethylsulfonyl)Imide [Hbet][Tf2N] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2723 G. Mawire and L. van Dyk Development of Scandium-Recovery Process from Titanium-Smelting Residue . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2735 Kota Nakashima, Yoshifumi Abe, Hidenori Okamoto, Akira Yoshimura, Matsuhide Horikawa and Seiichiro Tani Early Separation of Cerium from Mixed Rare Earths: A Review of Methods and Preliminary Economic Analysis . . . . . . . . . . . . . . . . . . 2743 J. R. Goode Oxidative Removal of Cerium from Rare Earth Elements Mixed Chloride Solution . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2753 Maziar E. Sauber Alternatives to 2-Ethylhexyl Phosphonic Acid, Mono-2-Ethylhexyl Ester for the Separation of Rare Earths . . . . . . . . . . . . . . . . . . . . . . . . . 2765 James E. Quinn, Karin Soldenhoff and Geoffrey W. Stevens Production of Partially Separated Rare Earth Elements (REE) from a Quebec Deposit . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2777 Jean-François Boulanger, Keven Turgeon, Claude Bazin, François-Olivier Verret and Dominic Downey Separation and Purification of Rare-Earth Elements Based on Electrophoretic Migration (PART II) . . . . . . . . . . . . . . . . . . . . . . . . 2797 P. Hajiani Radionuclide Removal from Ore and REE-Bearing Mineral by Leaching and Ion Exchange Separation . . . . . . . . . . . . . . . . . . . . . . 2807 Nicolas Reynier, Laurence Whitty-Léveillé, Cheryl Laviolette, Maxime Courchesne, Jean-Francois Fiset and Janice Zinck Influence of Substrate Properties on the Selective Leaching Performance of Cobalt from Cemented Carbides . . . . . . . . . . . . . . . . . 2829 Gregor Kücher, Stefan Luidold, Christoph Czettl and Christian Storf Rare-Earth Elements Recovery from Nd-Fe-B Hard Magnets by Hydrometallurgical Processes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2837 Haydar Güneş, Hüseyin Eren Obuz, Ezgi Oğur, Furkan Çapraz and Murat Alkan

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Sulfide Flotation

The Value of Incremental Performance Improvement in Concentrators—How to Secure and Quantify Small Gains . . . . . . . . . . 2847 Norman O. Lotter and Tim J. Napier-Munn The Use of Diagnostic Leaching for Flotation Insight . . . . . . . . . . . . . . 2859 Kymberley Worrell Molybdenite Polytypism and Its Implications for Processing and Recovery: A Geometallurgical-Based Case Study from the Bingham Canyon Mine, Utah . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2869 Craig R. McClung The Importance of Understanding Mineralogy Drivers for Flotation Performance . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2881 Zachery Zanetell and Jennifer Thogerson The Formulation and Use of Mixed Collectors in Sulphide Flotation—Valuable Performance Gains . . . . . . . . . . . . . . . . . . . . . . . . 2889 Norman O. Lotter and Deidre J. Bradshaw The Role of Soluble Sodium Silicate for Enhancing Flotation Selectivity of Sulphides Towards Grade and Recovery Improvements: Example from a Cu-Sulphide Ore . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2901 L. Xia, B. Hart, V. Sidorkiewicz and D. Shaw Froth Pumping Using Warman® Centrifugal Slurry Froth Pumps . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2915 Pavol Loderer and Aleks Roudnev A Fundamental Study of Disodium Carboxymethyl Trithiocarbonate (Orfom®D8) in Flotation Separation of Copper-Molybdenum Sulfides . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2927 Simon Timbillah, Courtney Young and Avimanyu Das Author Index . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2947 Subject Index. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2955

Part I

Pyrometallurgy Keynotes

Role of Research in Non-ferrous Metallurgy Development—Peter Hayes’ Contributions to Modern Pyrometallurgy Phillip J. Mackey and Evgueni Jak

Abstract A whole new group of non-ferrous pyrometallurgical processes with the ability to utilize tonnage oxygen and capture sulphur dioxide were developed and commercialized over the period 1940s–1980s. Many of these earlier process developments were undertaken with limited knowledge of the process chemistry and influence of key process variables; in many cases, piloting helped provide much new physico-chemical data, but gaps remained. The second generation versions of these technologies of today provide all the primary copper, nickel and lead produced worldwide by pyro-metallurgical smelting. Further, process development has continued and a new generation of copper and lead smelting technologies have also been developed in China since the 1990s. The older reverberatory and blast furnaces have been progressively replaced by the newer technologies—a good example being the introduction in 1992 of the copper IsaSmelt technology at the Mt. Isa smelter in Australia where the former fluid bed roaster and reverberatory furnaces were replaced by a single new smelting unit, together with an acid plant for sulphur dioxide collection. The development of these new technologies was made possible by investment in fundamental and applied research. The lesson for the future is, in order to sustain these improvements, continued investment in research and development capability is required—to do otherwise is to risk obsolescence and lack of competiveness in the world market. Dr. Peter Hayes at The University of Queensland is one of the many researchers and process engineers who have contributed to the fundamental understanding of metallurgical processes over this period of rapid change in technologies. The present paper briefly outlines some of the many contributions Peter Hayes has made to the understanding of kinetics, mechanisms and phase equilibria in metallurgical systems, and metallurgical process development.

P. J. Mackey (✉) P. J. Mackey Technology Inc, Kirkland, QC H9J 1P7, Canada e-mail: [email protected] E. Jak PYROSEARCH, Pyrometallurgy Innovation Centre, The University of Queensland, Brisbane, Australia © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_1

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Keywords Peter Hayes Copper smelting Lead smelting Nickel sulphide smelting Nickel laterite smelting Thermodynamics High temperature experimental techniques



Introduction The availability of metallurgical chemical and physico-chemical data often in the past has lagged behind new process technology developments. The classic cases include Bessemer’s development of the pneumatic converter for steelmaking unveiled in 1856 before understanding of the control of phosphorus levels in steel was mastered, or some decades later, commercialization of the Manhes copper converter in France in 1880 occurred before data were available on copper levels in oxidized converter slags. In other words, often a process was typically developed independently of underlying knowledge. Closer to the present era, this was the situation regarding a number of the new copper, lead and nickel smelting processes developed over the period 1940s–1970s before theoretical and experimental physico-chemical data applicable to the new processes were available. In many cases, the new metallurgical systems and their control were not only different but were considerably more complex than those of the older processes. For example, when the Outokumpu flash furnace was first piloted in Finland in 1947, much physico chemical data were lacking—in effect, original data were generated as piloting progressed. Understanding the phases present in the Noranda Process in the 1970s represented a stretch of known metallurgy that had largely focused on the fuel-fired reverberatory furnace—the workhorse technology of the industry for over a century. The Mitsubishi process, commercialized in 1974, utilized a new slag system for nonferrous metallurgy referred to as a “calcium ferrite” slag based on the CaO–FeO–Fe2O3 system. While generally known in the iron and steelmaking literature at that time, it still needed considerable work for application to the conditions of the new copper smelting system. The system was the subject of significant test work in Japan in the 1960s– 1970s. It is noted that the late Professor Akira Yazawa and colleagues at Tohoko University in Japan and the late Professor Reinhardt Schuhmann of MIT and later Purdue University, USA had both pioneered improved understanding of the thermodynamics of the then conventional copper smelting systems from about the 1950s. Thus, initially there were gaps in the literature or large variations in applicable thermodynamic data ranging from information on the free energy of formation of oxides for example, to activity data of certain solutes in liquid metal or slag phases. Peter Hayes is considered as one of the trail blazers in both developing and employing techniques for experimentally identifying reaction mechanisms and measuring chemical equilibria in pyrometallurgical systems, and expanding metallurgical knowledge in these areas. This paper explores Dr. Hayes’ contribution to this story of metallurgical development. It is noted that Peter Hayes’ work extended

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into many aspects of ferrous and non-ferrous metallurgy for both primary and secondary production of metals, however the present paper deals with one corner of Peter Hayes’ work—nonferrous smelting of copper, lead and nickel. Dr. Hayes work in other systems is acknowledged and a full list of his published papers is included in a separate paper by the present authors in this Symposium volume.

Non-ferrous Process Developments During the period before the 1940s up to about the 1980s, most of the world’s primary copper, nickel and lead was produced using reverberatory and/or blast furnace technologies plus the Peirce-Smith converter. A case in point was the Mt. Isa smelter in Australia which used the reverberatory furnace up until 1992, or the present Glencore smelter in Sudbury, Ontario, Canada, which used the nickel blast furnace up until about 1980. It is not a coincidence therefore that much of the early experimental work on phase equilibria and thermodynamics pertaining to copper smelting explored matte and slag systems common to reverberatory/Peirce-Smith converter systems. Examples are seen in the early work by Schuhmann et al. [1], Korakas [2] and Taylor et al. [3] to just name three (although the latter two studies extended measurements to the range of oxygen potential also applicable now to high-grade matte smelting and continuous converting operations). It was known, for example, that CaO in reverberatory furnace slag gave lower copper losses, however the exact mechanism and relationships were not known (for example, slag at the Gaspe smelter in Canada with 7% CaO (and 39% SiO2) assayed about 0.27% Cu (for 33% Cu matte). This result at the Gaspe smelter was the lowest copper level in slag out of 26 smelters [4]). At the time of the original development of the Noranda Process (for example the period mid-1960s to mid-1970s), there were insufficient physico-chemical data available to fully describe the new slag-matte-metal system involved; laboratory and pilot plant data thus provided much of the early information. One of the present authors (PJM) recalls that both Ruddle’s classic text [5], first published in 1953, and the text, “Phase Diagrams for Ceramists” published in 1964 [6] were constant reference books at Noranda. The FeO–Fe2O3–SiO2 phase diagram (Fig. 1) provided important liquidus data at the time [6]; however, the impact of oxides such as Al2O3, CaO, MgO or ZnO at Fe3O4 saturation was not well known and had to be extrapolated from available diagrams in Ref. [6]—which were in fact typically based on experimental work under conditions of iron-saturation and hence not readily applicable. Published information on the behaviour of minor elements in copper smelting at that time was not generally available. In one study of the behaviour of minor elements in the Noranda Process [7], engineers were forced to turn to early distribution data from publications in Transactions of the AIME dating to the 1900s!

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Fig. 1 The FeO–Fe2O3–SiO2 system [6]. This diagram was important in early development work on copper smelting processes before details of the impact of other oxides were known, such as those provided by Peter Hayes and co-workers

Gradually the picture changed with a large body of research work published on the emerging new copper smelting processes by the 1980s–1990s (for example refer Ref. [4]). Pyrometallurgy continues to play the dominant role in primary metal production from copper, lead, nickel and complex bulk lead/zinc sulphide concentrates, and also nickel from laterites. Increasingly, however, secondary (recycled) copper and industrial waste materials are being incorporated into existing primary metal

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operations and in addition, new dedicated pyrometallurgical reprocessing facilities are being designed and implemented. The industry faces ongoing challenges as the characteristics of these waste and recycled materials have changed with time. Lead-containing cathode ray tubes, once a major hazardous waste, are no longer used in televisions and stocks of used appliances of this type are dwindling, replaced by liquid crystal and plasma display devices. Lead acid starter batteries in vehicles will be increasingly replaced in the future with batteries having very different components for use in powering electric cars. Copper wire telephone landlines are being replaced by mobile phones, increased quantities of waste electric and electronic equipment (referred to as WEEE) containing a wide range of valuable metallic elements are being collected and treated. These technological changes cannot be efficiently and effectively undertaken without corresponding advances in fundamental understanding of the underlying process metallurgy and the development of tools to assist in quantifying the impacts of process changes. These tools and devices we now take for granted, but perhaps we do not appreciate that they were not always there as further discussed below.

Recent Advances in Metallurgical Research Capabilities There have been major advances in our knowledge and understanding of process fundamentals, and research capabilities over the past 50 years, from a time before the digital age and the use of modern computer tools. • Thermodynamic calculations have progressed in capability from regular solution models to the sophisticated, multi-component, multi-phase computer-based predictive platforms of today. These enable routine calculation and prediction of the outcomes of chemical reactions, replacing simpler approaches, using software packages such as Factsage. • Thermodynamic data have been extended from phase diagrams in binary, and selected ternary and higher order systems, to complex, multi-component systems in alloys, slags, mattes and solid solutions. Laboratory data on elemental distributions between phases have also been extended to more complex systems, approaching those encountered in industrial plants. • Heat and mass transfer calculations are no longer confined to analytical mathematical solutions of thermally activated processes, describing mass and thermal diffusion-limited processes for simple geometries. The experimental characterisation of fluid flow phenomena by using cold physical models as analogues for high temperature systems has been largely superseded by the use of computational fluid dynamic (CFD) tools, enabling decreased design and development costs, and the optimisation of reactor geometries; still, however, measurements such as tracer tests have a place in verifying model work.

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• Bulk chemical analysis and optical metallographic techniques can now be supplemented with detailed examination of sample microstructures using electron microscopy (scanning electron microscopy or SEM). Accurate measurement of the compositions of individual phases is now possible using electron microanalysis techniques, such as, energy dispersive (EDS) and wavelength dispersive spectrometers (EPMA). The presence of defects in crystal lattices and even the positions of individual atoms can now be resolved using high resolution transmission electron microscopy (TEM) techniques. This enables the forensic analysis of samples to be undertaken, providing invaluable insights into the conditions present in commercial furnaces—conditions that could not previously be characterised or measured. • Spectrometric techniques provide identification and quantification of ionisation states of elemental species in solutions. • Online measurement of solution composition, the effective oxygen potentials in metallurgical phases and gas analysis can assist in process control. • Older batch processes are being progressively replaced by continuous process technologies, e.g. new flash and bath continuous converting technologies, or continuous lead and tin refining techniques, providing increased energy efficiencies and reduced environmental impact. • Computational tools enable mass and energy balances to be undertaken on complex process flowsheets enabling improved design, reduced capital costs and process optimisation. Where did these advances, along with necessary support for the industry, come from? These outcomes are largely the result of fundamental and applied research undertaken in our Universities, government sponsored research laboratories and selected companies. Fundamental and applied research has contributed in a number of ways. Clearly, this research has resulted in the building of fundamental knowledge and the measurement of phase property data, and the development of theoretical models and research and development tools. These activities also provided the opportunities to generate develop and incubate new process ideas and concepts, step change/ disruptive technologies. This research has not only provided deeper and broader scientific understanding, but also research training, and the opportunities to transfer specialist knowledge and skills to the bright minds needed by industry today. The reality is then that industry today has benefited from the investment of time, effort and resources made by previous generations. We are all aware it takes time, years and in some cases decades, to develop new process concepts and translate these ideas into industrial practice. Strategic thinking and commitment is required to ensure that the pipeline of ideas and technical improvement does not dry up due to lack of investment in people and ideas. We need to invest in the future in order to survive in the competitive and changing world that is the metallurgical industry.

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9

We need to continue to ensure that we, as an industry, support such efforts, and we have access to and fully utilise the research capabilities needed to both continue efficient operations and to support change.

Peter Hayes’ Contributions to Metallurgical Research and Support for Metallurgical Process Development As pointed out in the introduction, Peter Hayes and his colleagues at The University of Queensland have developed and employed new experimental techniques for the characterisation of pyrometallurgical systems, extending the knowledge of process metallurgy fundamentals. Peter Hayes’ research on the mechanisms and kinetics of heterogeneous solid/ solid, gas/solid and liquid/solid reactions started with studies on heterogeneous reactions taking place during the nitriding of steels. As part of this work, experiments were devised to provide the first direct evidence for diffusion limited and chemical reaction interface control of precipitate growth in solids through the use of in situ transmission electron microscopy [8]. This interest in heterogeneous reaction mechanisms and the use of electron microscopy as a key tool to examine metallurgical systems was later extended to detailed studies of the mechanisms of gaseous reduction of iron oxides. Figures 2a–d illustrate examples of some the different interface structures that have been identified to date in solid-gas systems [9–11]. Peter Hayes has demonstrated that these heterogeneous gas/solid reaction systems are part of a broader class of analogous fluid/solid reaction systems, including the solidification of melts systems, whose behaviour are governed by instability criteria at moving interfaces [12]. These criteria determine the transition from planar to cellular and dendritic product morphologies [13]. These were not just academic studies—their application extended to practical systems. The example shown in Fig. 3 is the first known reported observation of nano-sized Ni–Fe alloy particles formed as a result of gaseous reduction of nickelcontaining saprolite ore in the Caron nickel production process, in use at the former Queensland Nickel operation in Queensland. Traditionally the Caron process was used on limonitic ores, this investigation explored extending this to saprolitic laterite ores. Counterintuitively the overall recovery of nickel in subsequent ammonia leaching was observed to decrease on increasing the reduction temperature. This was shown [14, 15] to be the result of the encapsulation of the alloy particles in a recrystallized olivine matrix (a known but little understood phenomena), which was impermeable to the ammonia leach solution. As told to the present authors, it was during a summer sabbatical at the laboratories of CSIRO in Melbourne, Victoria, Australia in the mid-1980s that Peter Hayes recognised the gap in fundamental slag equilibria data needed to understand and optimise the operation of the new technologies. Dr. Hayes was working

10

P. J. Mackey and E. Jak

Fig. 2 a–d—Examples of different interface structures formed during gaseous reduction of metal oxides in H2/H2O gas mixtures, a Continuous coupled growth of gas pores and iron metal, b Instability formation in wustite resulting in periodic splitting of the wustite product layer and discontinuous growth, c Continuous growth of gas pores covered with dense iron, d Bursting of dense nickel layers on NiO leading to discontinuous growth [9–11]

alongside Dr. W. T (Bill) Denholm who was carrying out test work on the then new Sirosmelt process for lead and copper smelting, pioneered earlier by Dr. John Floyd of CSIRO [16]. This realisation prompted the start of pioneering work on the development of new techniques for phase equilibria characterisation and the start of a long-standing research partnership with Evgueni Jak in the early 1990s, which continues today. New techniques for studying gas-slag-solids and gas-slag-matte-solids equilibria were developed using electron microprobe X-ray microanalysis with Wavelength Dispersive Detectors enabling accurate measurement of the individual phases present in the samples to be made rather than relying on the bulk mixture composition and X-ray powder diffraction to characterise the system [17, 18]. Examples of this research include, the first descriptions of the liquidus of lead IsaSmelt and other high lead/zinc smelting slags by the pseudo-ternary system ZnO–“Fe2O3”–(PbO+CaO+SiO2) in air, as illustrated in Fig. 4 [19, 20]. This was followed by systematic studies of the liquidus of ZnO–Fe2O3–(CaO+SiO2) slags at iron metal saturation [20], relevant to zinc fuming, and closely related to lead and zinc blast furnace slags. The liquidus and isotherms in the Cu2O–Fe2O3–SiO2

Role of Research in Non-ferrous Metallurgy Development …

(a)

11

(b)

Reduced at 800oC

2

4

3

1

20 nm

(c) -Above four patterns Fig. 3 a TEM image of nano-sized Ni–Fe alloy particles formed in nickel-containing serpentine reduced in 15%H2/N2 for 30 min at 800 °C; b Selected area diffraction (SAD) pattern ((010) FCC) of Ni–Fe alloy; c Energy dispersive spectrometer (EDS) spectra of points shown in Fig. 2a [7, 8]

system in equilibrium with metallic copper at high copper oxide concentrations is shown in Fig. 5 [21, 22]—directly relevant to direct-to-blister copper smelting and conventional copper converting, including continuous converting. Of interest, data from Ruddle et al. [3] noted above were seen to be favourably shown in the figure. The effect of the ratio SiO2/Fe and the ratio CaO/SiO2 on the measured liquidus temperature at fixed oxygen pressure is illustrated in Fig. 6 for the ferrous calcium silicate slag system; liquidus measurements on the same system at copper metal saturation have also been undertaken [23, 24]. The application of the techniques has been extended to non-silicate slags systems, such as the liquidus of the Cu2O–FeO– Fe2O3–CaO system at copper metal saturation relevant to Mitsubishi converter, KUCC flash converter and IsaConvert TSL copper converting technologies using Ca fluxing [25]. Slags systems describing the ferronickel smelting of nickel laterites have also been characterised through experimentally measured liquidus isotherms in a high MgO slags in the system MgO–FeO–SiO2–Al2O3 [26].

12

P. J. Mackey and E. Jak

Fig. 4 Experimentally determined liquidus in the system ZnO–“Fe2O3”–(PbO+CaO+SiO2) system in air; CaO/SiO2 = 0.6, PbO/(CaO/SiO2) = 4.3) [19]

Important work has been carried out on providing a clearer understanding of the formation of protective freeze linings on furnace walls that have externally-mounted copper coolers [27]. In recent years the techniques have been further improved and it is now possible to accurately measure chemical equilibria in multi-component, multi-phase systems extending the range of compositions to non-oxide systems, for example gas/slag/ matte/metal/speiss/solid systems [28]. Minor element distributions between these phases can also be measured. Examples of these latest developments are to be found in the present volume [29]. Realising early on the potential of this new experimental approach and the value of the new data, Peter Hayes established strong working relationships with Professors Art Pelton and Chris Bale and the Centre for Research in Computational Thermochemistry (CRCT) at Ecole Polytechnique in Montreal, Quebec, Canada. Starting from 1995, Evgueni Jak whilst working on projects related to coal ash slags, spent a year at CRCT and subsequently revised the (Al–Ca–Fe–Si–O) and

Role of Research in Non-ferrous Metallurgy Development …

13

Fig. 5 Liquidus isotherms in the “Cu2O”–“Fe2O3”–SiO2 system in equilibrium with metallic copper-iron alloy (data points in the pseudo-binary “Cu2O”–SiO2 in equilibrium with metallic copper [21] and in the pseudo-binary “Fe2O3”–SiO2 in equilibrium with metallic iron. pO2 in atm.) [22]

Fig. 6 Experimental measurements defining the effect of temperature on the liquidus of the “FeO”–CaO–SiO2 system in terms of CaO/SiO2 and SiO2/Fe weight ratio at an oxygen partial pressure of 10−6 atm. [23, 24]

14

P. J. Mackey and E. Jak

developed lead-zinc and coal (Al–Ca–Fe–Na–Pb–Si–O–Zn) databases for FactSage. This initiative has resulted in more than a 25-year long collaboration between the research teams in PYROSEARCH and CRCT, in particular Sergei Degterov, and support of the development of FactSage thermodynamic oxide databases [30, 31]. A chronology of selected major research initiatives by Peter Hayes in relation to process developments is of interest, since his career paralleled many key industrial developments. Figure 7 illustrates the chronology of selected major research initiatives by Peter Hayes, Evgueni Jak and coworkers in copper and lead smelting as shown on the x-axis in relation to the year of commercialization of the major copper and lead smelting technologies in use today (y-axis). Each bar on the x-axis indicates a specific initiative, with the bar height denoting year of commencing the work. (Data here are approximate.). It is seen that this research was at the centre of the major developments then underway, and importantly, contributed to the process optimization activities that followed. For reference, details on the chronology of the development of the major commercial technologies in copper, lead and nickel smelting are presented in Table 1.

2020 2010 2000 1990 1980 1970

Two-Step (2015) SKS Conv (2013)

SKS smelƟng (2001) Noranda Conv. (1998) FCF (1995) Isasmelt Cu (1992), Pb (1990) QSL (1990)

Kivcet(1981) OKO Dir Cu (1974 Mit., El Ten(1974) Noranda (1973) Outk O2(1972) Inco FF (1952) Outk-FF(1949)

1960

Fig. 7 The chronology of selected major research initiatives in lead and copper smelting by Dr. P. Hayes and co-workers shown on the x-axis in relation to the year of commercialization of the major copper and lead smelting technologies in use today. Each bar on the x-axis indicates a specific initiative, with the bar height denoting year of commencing the work. Data are approximate

Role of Research in Non-ferrous Metallurgy Development …

15

Table 1 Chronology of the development of major commercial technologies in copper, lead and nickel smelting Process Non-ferrous electric furnace smelting

Year started Laboratory tests

Pilot tests

Commercial plant

∼1917

∼1917

1917

Electric furnace-PGM concentrates

No lab tests

Year not known

1969

(Outokumpu) flash furnace (air)

No lab testsa

1947–1948

1949 (O2 1972)

Inco flash furnace

1945–1946

1946–1949

1952

Vanuykov process (Cu, Ni)

∼1950s

1950–1970s

1977

Outokumpu direct Cu

1960s

1969

1978

Fluidized bed- Electric furnaceb (Ni)

NA

Used Coy. data

1960

RKEF Fe–Ni smelting

NA

Likely early 1950s

1955

Noranda process

1963–1968

1968–1972

1973 (O2 1976)

El Teniente process

NA

1972–1974

1974

Mitsubishi process

1961

∼Mid 1960s–1972

1974

Ausmelt process (Pb, Sn, Cu, Ni)

1971

1981–1992

1992

Isasmelt process (Pb, Cu)

1978–1979 (Pb)

1980–1985 (Pb), 1987 (Cu)

1990 (Pb), 1992 (Cu)

QSL (Pb)

Late 1970s

1981

1990

Kivcet (Pb)

∼1950s

1950s

1984

Kennecott-Outokumpu flash converter

Mid-1980s

1984

1995

Noranda converter

No lab tests

1996–1997

1998

SKS/BBM smelting

NA

1990–2000

2001

SKS converter

NA

200% to 11,320 kt/y, respectively [5]. Giving this explosive growth in base metal demand, it seems reasonable trying to understand, which are the challenges that the base metal industry will face, and in particular, where sulfide smelting opportunities and needs will emerge. The present work will try to give an overview on the copper industry in aspect related to: • • • • •

Copper concentrate supply Development of thermochemical knowledge to support process understanding Digitalization Technology development Process fundamentals and value integration

Table 1 Worldwide copper production comparison between 1988 and 2017 [2, 3] Item

1988

2017

Estimated concentrate production, tonnes Copper production from primary smelting, tonnes Estimated slag production, tonnes Worldwide estimated sulphur capture,% Estimated acid production, tonnes

28,450,000 8,537,000 14,300,000 45 12,500,000

59,300,000 16,050,000 29,600,000 90 54,300,000

42

G.R.F. Alvear Flores et al.

% of Copper Production

(a)

16

Top World Copper Smelting Countries in 1988 (Total Production: 8.25 Mt/y) Including Secondary Production

14 12 10 8 6 4 2 0 Chile

% of Copper Production

(b)

USA

Japan

USSR

Canada

China

Poland

Country 40

Top World Copper Smelting Countries in 2017 (Total Production: 19.06 Mt/y) Including Secondary Production

35 30 25 20 15 10 5 0 China

Japan

Chile

Russian FederaƟon

Zambia

South Korea Germany

Country

Fig. 1 a Top world smelting countries in 1988, b Top world smelting countries 2017

Understanding the Supply In parallel with the increasing demand for copper, concentrate supply has gradually changed. Copper smelters have noticed a reduction in copper content in concentrates associated with a continuous decrease in ore grades and increasing ore complexity. As existing open pit copper mines move deeper into primary mineralisation with higher proportion of chalcopyrite and bornite and most of new mines seems to be based on deeper primary mineralization, copper smelters should expect lower concentrates grades. These concentrates will have lower Cu/S and Cu/Fe ratios, higher gangue contents, higher impurity contents and finer granulometry resulting in an increase in slag and sulphuric acid productions, potential higher dust generation (uneven mineralogy distribution for each size fraction) and therefore less units of copper per unit of processed concentrate. An additional important factor is the increase in impurity contents in copper concentrates. Table 2 shows a comparison for the contained As, Bi, Pb, Hg and F content in copper concentrates for 2012 and 2016. The analysis is based on Wood Mackenzie concentrate assays with a capture of 75–78% of Cu production for 2012 and 88% for 2016 [7].

Sulfide Smelting: Thirty-Five Years of Continuous Efforts …

43

The increment in arsenic concentration is an example of the gradual additional pressure exerted on smelter performances, as they need to find solutions to optimize their global impurity management and integrate adequate technical solutions to stabilize waste streams. Under this scenario, it seems reasonable to understand how a decrease of copper grade will affect the competitiveness of existing smelters and what cost-effective measures shall each one of them take into consideration to remain competitive; in other words, how they will differentiate from each other. Clearly adequate technology selection and performance will play a crucial role in this race to enhance (i) superior metallurgical performance, (ii) a sustainable and environmentally friendly operation, (iii) adequate impurity management and (iv) proper product quality. However, a technology change for most of smelters is not an option, as large CapEx cannot be justified by “just the need” to become adequate to new supply. Asset optimization becomes then necessary. Key aspects to adjust operations to the change in supply will be: • Flexibility of the process configuration to adapt to new grades: Find new operating conditions and adjust production to this reality. • Interaction with concentrate suppliers and ability to blend and stabilize their smelting asset on regular basis if possible. • Ability to maximize their respective impurity capacity. • Higher copper prizes challenging current technologies ability to maximise copper recovery. • Slower copper demand. • Energy utilization: how to adjust operating conditions for a given feed to optimize energy efficiency and utilization. • Ultimately, considerations for new/ improved technologies to meet the challenge of resource efficiency. Coursol et al. tried to quantify the above-mentioned challenges in an indirect way by quantifying the specific energy consumption associated to different technology configurations [5]. However, their study applied similar feed compositions to each technology configuration in an effort to equalize the input conditions and generate

Table 2 Comparison of contained As, Bi, Pb, Hg and F in copper concentrates between 2012 and 2016 [7]

As Bi Pb Hg F

t t t t t

Total content in Cu concentrate 2012 2016 Δ (%)

Average grade 2012 2016

54,600 3,820 82,000 107.3 3,837

0.15% 0.012% 0.29% 3.4 g/t 121 g/t

107,300 5,510 106,000 213.6 7,517

97 44 29 99 96

0.23% 0.013% 0.28% 5.6 g/t 182 g/t

44

G.R.F. Alvear Flores et al.

comparable trends. This methodology may need to be slightly modified to compare technology configurations based on their typical supply, as technology are chosen to meet particular feed characteristics.

Integration of Fundamental Knowledge into Smelter Performance Knowledge integration plays a key role in the conceptualization and development of metallurgical processes. This has been demonstrated in the last fifty years as metallurgical processes such as flash and bath smelting technologies have been developed and following up methodical approaches to generate fundamental, technological and operational knowledge. In most cases, these technologies have been developed to meet specific challenges associated to energy demands, environmental constrains or particular materials requiring processing. As these conditions change, technologies are forced to deliver results beyond their originally conceived technological borders. Metallurgists are therefore required to think outside the box and deliver operational solutions that meet management expectations. Factors such as increase of minor element concentrations in concentrates, reduction in Cu/S ratio, changes in mineralogical composition and increase in minor oxide concentrations have to be considered to define fundamental and operational aspects that will enable a safe and stable operation of a particular metallurgical asset. Integration of experimental and fundamental approaches are required to provide a holistic solution that combines metallurgical fundamental knowledge, process design considerations, engineering improvement aspects and smart use of data.

The Development of Computational Modelling Since pyrometallurgical pilot tests work is very expensive, it should be reduced as much as possible. Also, with ever-increasing complexity of the feed mix there is a higher demand to model process metallurgies to understand optimization potential more effectively. Due to that, chemical thermodynamic knowledge is of great help for the optimization of the pyrometallurgical processes; this approach potentially can lead to improving the economy of plants and the ecology of planet. In 1970s and 1980s, valuable fundamental chemistry data (such as sulfur-oxygen potential diagram for Cu–Fe–S–O–SiO2 system, distribution of minor and major elements etc.) was generated for the smelting and converting of the copper making processes [8–18]. These modelling efforts were instrumental to support development of key new technologies such as the Mitsubishi Process, Noranda process and Sirosmelt and growth the knowledge of consolidated process such as the Flash smelting process.

Sulfide Smelting: Thirty-Five Years of Continuous Efforts …

45

That approach for the full characterization of the today’s pyrometallurgical process is considered as the semi-quantitative method. Furthermore, mathematical models were developed by Professor Sohn and his team to thermodynamically describe the partitioning of minor element such as Bi, As, Sb, Zn, and Pb between gas, slag, matte and copper metal in smelting, converting and slag cleaning processes [18]. Nagamori and Mackey developed a computer-based thermodynamic methodology to simulate the smelting and converting process of Noranda to describe the partitioning of major and minor element for n major elements of Cu– Fe–O–S–Si [19, 20]. From early 2000, our capability in computational modelling of pyrometallurgical processes has gone through sharp rise and there are a number of software packages available for advanced thermodynamic modelling. Generally these models use solution models to calculate the non-ideal behaviour of metallurgical melts such as slag, matte, alloys as well as oxide solid solutions such as spinel, olivine etc. These models give accurate representation of experimental data on binary, ternary and quaternary systems. Most of these models have good capability in predicting the behaviour of higher order and more complex systems that one encounters in metallurgical processes. In some cases, the capability and limitations of these models have been assessed through validation against targeted measurements on multi-component melts and solid solutions. Due to a significant rise in the complexity of concentrates for primary productions and the recycling feed stream, from late-1990 s, commercial thermodynamic packages such as CSIRO MPE, MTDATA and FactSage were developed and optimized to calculate the partitioning of major and minor elements between phases. For example, distribution of minor elements (i.e. As, Sb, Bi, Pb, Zn, Se, Te, Cr, Sn) between slag, matte and liquid copper were included in the thermodynamic database developed by CSIRO MPE [22–24]. In addition to the phase equilibria information, the MPE package offers the prediction of slag viscosity in multi-component and multi-phase system [24]. The modified quasichemical model was used in FactSage to develop the thermodynamic properties of the slag, matte, and liquid copper phases in the Cu–Fe–Si–O–S–Ca system [25]. Another example is an ongoing research program implemented by PYROSEARCH in the University of Queensland to characterize the multi component copper containing system (Cu–Fe–O–S–Si–Al–Ca–Mg) and multi-phase (gas/slag/ matte/metal/solids) systems with minor elements (Pb, Zn, As, Bi, Sn, Sb, Ag, and Au). An integrated experimental phase equilibria and thermodynamic modelling approach (i.e. FactSage) has been undertaken [26]. The question is if the available databases can be held accountable for the industrial needs. In general, from development of thermodynamic databases to optimization of the existing ones, the critical assessments and evaluation of the available thermodynamic and phase equilibrium data are necessary. The review of available literature for slag and matte in the low-order system of Cu–Fe–S–O–Si indicates significant differences in the reported phase equilibrium data; the inconsistency between the experimental data reported by various researchers may lead to development or optimization of an uncertain database. For example, as can be seen

46

G.R.F. Alvear Flores et al.

from Fig. 2, predictions performed by MPE32 [24] and FactSage public database [25] for the dissolved copper in slag vs matte grade for the given conditions are much lower than the majorities of the previous research studies. The ability to correctly understand and predict equilibrium copper contents in slag as function of matte grade gives also opportunity to understand limits of current metallurgical processes and therefore define options to improve metal recoveries. This will have a strong economic impact on copper smelters. In addition to the role of uncertainties in the development of thermodynamic databases driven from fundamental experimental data, the kinetics plays an important role in interpretation of the industrial plant data as well. Importantly, the research studies mainly have been based on the measurement of the portioning of major and minor elements between the condensed phases. There is a great need to experimentally measure the accurate distribution of volatile species between gas and condensed phases.

Solutions on Furnace Integrity Using Fundamental Knowledge Ways of increasing productivity have been always a challenging issue for pyrometallurgical processes. There have been some improvement in the integrity of pyrometallurgical reactors using traditional approach with refractory lining or more novel and advanced solutions incorporating freeze lining.

100∙Cu/(Cu2O+FeO+SiO2+S) in slag (wt pct)

Fallah-Mehrjardi et al, P(SO2)=0.25 atm MPE32, P(SO2)=0.25 atm Tavera et al,1979, P(SO2)=1 atm Korakas, 1964, P(SO2)=0.1 atm

8

FactSage, UQPY database, P(SO2)=0.25 atm FactSage public database, P(SO2)=0.25 atm Korakas, 1964, P(SO2)=1 atm Henao et al, 2013, P(SO2)=0.1 atm

T =1250 °C Cu-Fe-S-O-Si gas/slag/matte/tridymite

7 6 5 4 3 2 1 0

10

20

30

40

50

60

70

80

100∙Cu/[Cu+S+Fe] in matte (wt pct) Fig. 2 Comparison of the previously reported data [27–29] and present results for the gas/slag/ matte/tridymite equilibrium in the Cu–Fe–O–S–Si system at 1250 °C, at controlled partial pressures of SO2 for copper in slag vs matte grade. Legend: data in the graph are from Fallah-Mehrjardi et al. [30], Korakas et al. [28], Tavera and Davenport [27], Henao and Jak [29], MPE32 [21], FactSage public database 6.2 [30], and FactSage UQPY database [31]

Sulfide Smelting: Thirty-Five Years of Continuous Efforts …

47

Refractory: The chemical dissolution of refractory into the slag occurs through the direct contact with aggressive slag and the infiltration of liquid into refractory via pores. The kinetics of chemical reactions between slag and refractory is improved due to high process temperature and forced convection in bath; this gives rise to the refractory corrosion, planned and/or premature relining repairs resulting in the increase of operating costs [32]. Most previous fundamental and applied research on refractory materials focused on the selection and manufacturing of suitable refractory materials. There are some fundamental research studies available in open literature to explain the refractory and slag interaction from chemistry point of view. In recent research studies [32– 34], a systematic seven-step approach for the slag/refractory interaction and the identification of the optimum slag/refractory combinations has been developed. From metallurgical point of view, the slag engineering approach for the slag/refractory interaction needs to be further investigated for the increase of reactor service life as well as design of the new processes. The effect of refractory dissolution on the chemistry of whole system needs to be assessed carefully and actions should be taken accordingly. Freeze Lining: Refractory materials can be replaced with a cooling system at which the slag layer (freeze lining) is solidified on the cooling walls of the reactor. This technology has been used in copper flash smelters, slag cleaning furnace, ilmenite smelting processes, zinc slag fuming and Hall–He´roult aluminum production [35]. Thick freeze-lining leads to decrease of the throughput while thin deposit gives rise to increase of heat loss and spalling resulting in damaging the cooling module. As a result, process conditions (such as bath temperature, flux addition, etc) should be adjusted in way to obtain optimum thickness of slag deposit [36]. In most research studies on freeze-lining, heat transfer modelling was the main focus of the investigation. A few studies highlight the effect of slag chemistry on the freeze-lining thickness and heat transfer characterization. These research studies indicated that effect of chemically related parameters of liquid slag (e.g., viscosity), fluid flow characteristics, and mass transfer rates at the stationary deposit interface, are additional key parameters that needs to be included in design considerations [35].

The Use of Big Data and Machine Learning for Operational Excellence To fit fundamental models to reality, the use of “Big Data” and machine learning algorithms is a profound step to predict process behavior. “Within the mining and metals industry, digitalization will be a force that changes the nature of companies and their interaction with employees, communities, government and the environment at every step of the value chain.”, according to the World Economic Forum’s Digital Transformation Initiative [37]. This digitalization trend, a manifestation of the “age of analytics”, is currently driving competition and disruption across many industries [38].

48

G.R.F. Alvear Flores et al.

Digitalization is defined as “the use of digital technologies to change a business model and provide new revenue and value-producing opportunities” [39]. In the metals industry, this includes the smart use of data and algorithms to improve complex operational decisions in real time. In practice, large amounts of historical sensor data, e.g. the operational history of running a smelter, are combined with the respective outcomes, the so-called prediction targets. In a second step, called Machine Learning, computer algorithms programmed by Data Scientist with support from domain experts extract knowledge from this data to predict future events. For example, a goal might be to reduce the amount of copper and precious metals remaining in the slag depending on the variables input concentrate compositions, amounts, temperature, and other relevant operating conditions. If there is a causal relationship between the variables and the target, and if the amount of data is sufficient to show this statistically (the more complex the problem, the more data one typically needs), the Machine Learning algorithms are able to learn the statistical correlations for driving operational excellence empirically. The ultimate goal is to implement the algorithms as real-time decision support and/or machine automation to generate business value. This approach can be used for a variety of operational business cases. The best practice is to start small and lean, i.e. to pick the use cases where large amounts of relevant data with sufficient quality are already present and a measurable business case promises quick and significant ROI (Return on Investment). A “minimum viable” proof of concept, using as little time and resources as possible, tests the viability and feasibility of the data science use case. If the ROI is proven, the results will be implemented in a small-scale pilot and then rolled out large scale. Many companies are demonstrating the commercial viability and feasibility of this data-driven Machine Learning approach in complex industrial settings. Google uses advanced Machine Learning algorithms to optimize the operations of its data centers, thereby reducing their cooling costs by 40% [40]. In the steel industry, Big River Steel’s newest steel plant contains some 50,000 sensors to collect the necessary data for Machine Learning [41]. This revolution is called “Industry 4.0”, and Machine Learning algorithms and capable Data Scientists are the enablers to extract knowledge and generate business value from these vast amounts of sensor data.

Technology Developments in Copper Industry Figure 3 is an attempt to capture some key technology innovations in the copper smelting processes over the last 30 years [42]. Over this period, the copper industry has witnessed a dramatic technology change as copper smelters have gradually modify their main smelting technologies to improve their operational performance, cost competitiveness and environmental performance. Technology providers experienced a golden age providing technology services around the world to support this change.

Sulfide Smelting: Thirty-Five Years of Continuous Efforts …

49

After the consolidation of the flash technology, bath-smelting technologies developed in Japan, Canada, Chile and Australia were developed aiming to bring innovative process solutions to deal with increasing energy cost and tighter environmental regulations. By 2015, there were not many copper smelters left behind to transform their respective operational configurations. Nevertheless, as feed complexity increases, we may continue witnessing evolutions to modify current processes searching for more flexible, environmentally friendly, cost effective and energy efficient technologies. A particular example in this context is the increasing effort done by Chinese technologists in the last decade to adapt existing processes and create new one to meet their country needs. In this sense, Copper 2013 held in Santiago became a milestone in the way western copper industry perceived and understood Chinese technology development in the copper smelting industry. This conference organized full sessions dedicated to understand the application of Bottom Blowing Technology to copper smelting, with participation of Chinese experts from the academic, engineering and industrial world. This was the first substantial interaction between Chinese and western copper technologists. Copper 2016, held in Kobe, followed the same example. However, despite this interaction, presented information still not enough to have a clear understanding of these developments. Available operational data stills not to the level required to generate clear an unbiased metallurgical and operational comparisons with current technologies used in western countries. Table 3 compares some metallurgical and technological aspects of the Bottom Blowing Smelting and Submerged Lance Smelting, two new expressions of bath smelting in China. Production of high matte grade (70%) at relatively high oxygen enrichment in the blast (74% O2) for lower grade materials are key features of both technologies. In addition, high Fe/SiO2 ratio slags (close to spinel saturation) and low copper content slag (3 wt%) are key characteristics helping them better first pass copper recovery compared with similar bath smelting technologies.

1973

Noranda Reactor Canada

1974

Mitsubishi Process Japan

1977

1978

Sirosmelt Lance Industrial test at MIM Australia

1985

Inco Flash Furance

1992

Cu ISASMELT™ USA & Australia

1995

1997

1999

2008

2015

Mega Smelters BBS

SLCR

2000 -

Noranda PT Smelting Kennecott Converter Mitsubishi Process Outokumpu Flash Converter

El Teniente Reactor Chile Direct to Blister, KGHM

Fig. 3 Some milestones in technology development for copper smelting industry since 1973

50

G.R.F. Alvear Flores et al.

Table 3 Comparison of selected parameters between bottom blowing technologies in China— SLS (Dongying Fangyuan) and BBS (ENFI) [44, 45] Parameter

Dongying Fangyuan submerged lance Smelting

ENFI bottom blowing smelting

Reactor

SLS

BBS

Matte grade

72–75

72–75

slag type

SiO2–FeOx

SiO2–FeOx

Fe/SiO2

1.8–2.0

1.8–2.0

Cu in Slag

1.7–2.0%

2–4.5%

Dust generation

2%

2%

Feed type

Continuous, wet from furnace top

Continuous, wet from furnace top

Feed rate

Actual: 156 tph concentrate Max 250 tph, total feed

Minimum: 70 tph Max 250 tph, total feed

Cu in feed, %

20–25% Cu

20–25% Cu

Enrichment, %O2

75% O2

up to 75% O2

Injectors

23 High pressure tuyeres, bottom two lines with different angle from bottom side tuyeres plus additional aide blowing at end of reactor

Relatively high Pressure, one line, 9 tuyeres in two lines as well.

Furnace size

5.5 m φ × 28.0 m

5.8 m φ × 30.0 m (largest furnace)

Temperature

1180 °C

1180–1200 °C

Dust Generation, % feed

1.7%

2%

A Pending Effort: Continuous Converting Continuous converting has been for several years a recurrent topic in the copper industry. Despite early adoption of the Flash Converting process by Kennecott, it took several years to have a second plant in operation. In 2009 at the TMS Annual Meeting in San Francisco, US, metallurgist celebrated the 100 years anniversary of the Peirce-Smith batch converting technology. A special session and workshop was dedicated to discuss the future of the technology and the likelihood of adopting continuous copper converting as standard technology. After almost ten years, the reality shows that batch converting is still the standard converting process for most of copper smelters despite tighter environmental regulations. Traditional solutions adding additional hooding system or enclosing converter aisles are still preferred, increasing overall cost and specific energy consumptions, limiting somehow flexibility. In China, however, metallurgists are moving faster towards continuous converting. The two main contributions are the Bottom Blowing Converting and the Submerged Lance Converting and Refining Technology. Some selected indicators of both technologies are shown in Table 4 [44–46]. From the metallurgical point of view, the SLCR process is a batch process as the iron and sulfur removals take gradually place increasing the copper concentration in the generated copper phase. However, the real breakthrough of the SLCR is the

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ability to combine in one reactor the batch converting and anode refining processes. A proper combination of oxygen, air, nitrogen and natural gas allows proper control of the oxygen potential. Most likely desulfurization of the blister copper can be achieved without increasing the oxygen content to levels normally observed in the blister generated in the bath copper converting. More operational information is required to produce a well-balanced metallurgical and operational assessment. However, it is clear that this development may change the traditional mindset of copper smelter operators. The BBC process on the other hand can be described as a continuous copper converting process that differentiates from the Ausmelt-Outotec, ISACONVERT and Flash Converting for their attempt to continuously feed molten matte. However, the heat balance of the furnace limits production capacity as cold matte is required to stabilize the heat balance of the process. Additional process data are required for better understanding of advantages and limitations of these two processes and quantify furnace campaigns.

Table 4 Comparison between the SLCR and the BBC processes [43–46] Indicator

SLCR process

BBC process

Feed type

Molten matte

Feed flow

Molten, continuous

Matte grade, %Cu Process type Feed rate Furnace Size Blowing gas Enrichment Heat Balance

72-75 Batch cycle(*) 55 t/h 4,8 m φ × 23.0.0 m Air-Oxygen-Nitrogen-Natural gas 21–30%O2 Addition rate not available but likely lower than BBC Process Variable through the process Variable through the cycle

Cold and hot matte to meet Heat Balance Solid and Molten, continuous 72-75 Continuous 20 t/h 4,1 m φ × 18.0 m Enriched air 38–40% O2 High dependency on cold reverts addition Fixed 38–40%

FeOx-SiO2-CuO0.5 1.20 Variable through the cycle >1180 °C

FeOx–SiO2–CuO0.5 1.20 1230 °C

10–12% 30 ppm Anode

8–12% 3,000–5,000 ppm Blister

Oxygen potential Oxygen Enrichment Slag Fe/SiO2 Process Temperature Cu in Slag %S in Cu blister Final Quality

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Towards Multi Metal Recovery: Positioning Aurubis at the Frontier in Non-ferrous Metallurgy With primary and secondary copper smelters operations Aurubis is one of the largest copper producers in the world. 4 shows some characteristics of the smelter operations of the company. Indicator

Hamburg Cu smelter Flash furnace Peirce-Smith converter anode Furnace

Hamburg Pb smelter Electric furnace Peirce-Smith converter

N.A.

Refinery

Electric Furnace Isa process

Product

Cathode

Configuration

Slag cleaning

Pb pyro refinery Lead ingot blister

Pirdop smelter Flash furnace electric furnace Peirce-Smith converter anode furnace Slag Flotation Isa process

Lünen Sec. smelter KRS+ (ISASMELT™, TBRC)

Olen Sec. smelter Contimelt

Slag Settling

N.A

Isa process

Isa process

Cathode, Anodes

Cathodes

Cathodes

The combination of primary and secondary copper and secondary lead smelters allows Aurubis to optimize process materials flows and minor metal production across its assets with the main aims of maximizing asset utilization, impurity capacities and value associated with the different supplies.

Getting a Grip on Complexity To optimize the operation of five metallurgical assets is a complex target itself. In addition to that, as complexity of materials increase, optimizing impurity processing capacities and value generated by these materials becomes even a more complex challenge. These days it is getting harder and harder to plan the raw material supply for complex production networks using traditional manual planning tools – and it is the same for Aurubis as well. At the same time, as the complexity of materials is increasing, the technical processability and the associated added value need to be determined even faster to be able to respond proactively in a dynamic market environment. The inherent high combinatorial complexity makes an Advanced Analytics approach valuable. Aurubis has developed an integrated approach to develop a value-based planning model. The model uses large scale mathematical optimization, an advanced analytics method for value optimization.

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The entire primary copper supply chain – from concentrate and secondary raw materials to cathode is mapped taking into account the technical restrictions and commercial constraints to enable a multi-site optimization of the raw material utilization. The model constitutes a first effort to combine elements of advanced analytics along with metallurgical knowledge to support the business. Although the chemical industry and the oil and gas industry have used advanced analytics for some time now, Aurubis is playing a pioneering role in the metalworking industry with this approach. The implementation of this integrated approach is helping Aurubis to improve resource efficiency across the different sites.

One level beyond Following the development and implementation of the optimization model for raw material supply, Aurubis has moved toward the next level to continue using mathematical optimization tools to support their operational planning and to optimize existing and future assets. Therefore, new efforts were made to develop value-based operational planning tools that covers the entire production process. The main driver was the potential of high economic and technical benefits resulting from implementing such an approach into the daily processes of a production plant. [47].

The Future: Multi Metal Recovery Copper is by nature a carrier for many other metals that are important for the society. Due to that, Aurubis steps even more into the direction of a multi metal producer, even if its evolution from a copper producer to a multi metal producer is not new. Through its history and development the synergies between copper in the primary and secondary production and lead production has allowed producing a number of valuable metals associated to copper and lead. In 2017, Aurubis has announced a new smelter project: Future Complex Metallurgy (FCM) that will contribute to enhance the multi metal recovery philosophy of the company. The project involves the development and implementation of an innovative metallurgical process to ultimately increase the operational result of the Aurubis Group. FCM will allow Aurubis to expand its raw material base with materials containing lead, sulphur and copper. This development makes the effort of a value-based multi-smelter network optimization by means of planning models even more attractive.

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Concluding Remarks In the last 40 years, the base metals industry has witnessed dramatic changes. Larger production increments, decrease in supply quality, technology development, increased energy efficiencies, environmental constrains, social awareness are among several aspects, key factors that today continue evolving. These factors will continue creating pressure on the industry and in particular in technologists to develop required new solutions that could meet evolving social standards. A holistic approach that gets through silos is required to bring required breakthroughs for the future. Dialogue between disciplines is also required. Effective metallurgical solutions successfully implemented in operating plants require not only fundamental chemical and metallurgical knowledge but also an integrated approach across the value chain that considers the needs from the supply (geology-mine-concentrator) and the product size (quality related to the particular applications) and society. Future supply will be a combination of traditional and urban mining, requiring the development of new process flowsheets that will combine the intrinsic properties of base metals to optimize the value. In particular, potential evolution of e-mobility will have a large impact on existing metal recovery flowsheets. Future process development will require integration of different disciplines. Metallurgist associated with mathematician, chemists, process control engineers, technology designers, mining engineers, environmental and energy specialist will have to work in a collaborative fashion to develop required process solutions. Aurubis is committed to embrace this future. For this purpose, combination of primary and secondary base metals metallurgy is an intrinsic advantage. However, a static view is no longer valid as new processes will require a constant transformation to meet the challenges imposed by a dynamic supply.

References 1. Sohn HY, George DB, Zunkel AD, Foreword, VI; Proceedings of the 1983 international sulfide smelting symposium and the 1983 extractive and process metallurgy meeting of the metallurgical society of AIME, San Francisco, California, November 6–9, 1983 2. International Copper Study Group, Internal Communication 3. WoodMackenzie, Global Copper Shorterm Outlook 2018 4. International Nickel Study Group, Internal Communication 5. International Lead Study Group, Internal Communication 6. Coursol P, Mackey PJ, Diaz CM (2010) Energy consumption in copper sulphide smelting. In: Proceedings of copper 2010, GDMB, pp 649–668 7. WoodMackenzie, Global Copper Shorterm Outlook (2016) 8. Yazawa A (1974) Thermodynamic considerations of copper smelting. Can Metall Q 13(3): 443–453 9. Yazawa A (1977) Trends in modern copper smelting processes. Erzmetall 30(11):511–517 10. Yazawa A (1980) Distribution of various elements between copper. Matte Slag Erzmetall 33 (7/8):377–382

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11. Yazawa A (1984) Slag-metal and slag-matte equilibria and their process implications. Metall Slags Fluxes Int Symp Proc 2:701–720 12. Acuna C, Yazawa A (1987) Behavior of arsenic, antimony and lead in phase equilibria among copper, matte and calcium or barium ferrite slag. Trans Jpn Inst Met 28(6):498–506 13. Itagaki K, Seino T, Yazawa A (1979) Thermodynamic activity of arsenic in the liquid copper-iron-sulfur mattes. Tohoku Daigaku Senko Seiren Kenkyusho Iho 35(1):1–10 14. Itagaki K, Yazawa A (1982) Thermodynamic evaluation of distribution behavior of arsenic in copper smelting. Trans Jpn Inst Met 23(12):759–767 15. Kashima M, Egouchi M, Yazawa A (1978) Distribution of impurities between crude copper, white metal and silica-saturated slag. Trans Jpn Inst Met 19(3):152–158 16. Nagamori M, Azakami T, Yazawa A (1989) Activities in the copper-iron-sulfur mattes at 1473 K. Metall Rev MMIJ 6(2):112–127 17. Nagamori M, Takeda Y, Yazawa A (1989) Statistical thermodynamics of ferrite slags. 1. Activities of oxides of calcium, iron, copper, cobalt, nickel, lead and zinc at 1523–1573 K. Metall Rev MMIJ 6(1):6–21 18. Kim H, Sohn H (1991) Computer analysis of minor element behaviour in copper smelting and converting under high oxygen enrichment, and in converting with calcium ferrite slag. Copper 91(Cobre 91) 19. Nagamori M, Mackey PJ (1978) Thermodynamics of copper matte converting: Part I. fundamentals of the Noranda process. Metall Trans B 9B(2):255–265 20. Nagamori M, Mackey PJ (1978) Thermodynamics of copper matte converting. Part II. Distribution of gold, silver, lead, zinc, nickel, selenium, tellurium, bismuth, antimony and arsenic. Metall Trans B 9B(4):567–579 21. Chen C, Zhang L, Jahanshahi S (2010) Thermodynamic modeling of arsenic in copper smelting processes. Metall Mater Trans B 41(6):1175–1185 22. Chen C, Zhang L, Jahanshahi S (2013) Application of MPE model to direct-to-blister flash smelting and deportment of minor elements. In: Proceeding of copper, Santiago, Chile 23. Chen C et al (2006) Thermodynamic modelling of minor elements in copper smelting processes. Minerals, Metals & Materials Society 24. Zhang L et al (2002) CSIRO’s multiphase reaction models and their industrial applications. JOM 54(11):51–56 25. Degterov SA, Pelton AD (1999) A thermodynamic database for copper smelting and converting. Metall Mater Trans B 30B(4):661–669 26. Jak E et al (2016) Integrated experimental phase equilibria and thermodynamic modelling studies for copper pyrometallurgy. In: 9th International copper conference, Kobe, Japan 27. Tavera FJ, Davenport WG (1979) Equilibrations of copper matte and fayalite slag under controlled partial pressures of sulfur dioxide. Metall Trans B 10B(2):237–241 28. Korakas N (1964) Etude thermodynamic de l’équilibre entre scories ferro-siliceuses et mattes de cuivre. Application aux problèmes posés par la formation de magnetite lors du traitement des minerais sulfurés de cuivre. PhD thesis, Univirsité de Liège 29. Henao HM, Jak E (2013) Experimental study of slag-matte equilibria in the Ca–Cu–Fe–O–S– Si system at fixed P(SO2) and P(O2). Private Communication: Pyrosearch, University of Queensland 30. Fallah-Mehrjardi A et al (2018) Experimental investigation of gas/slag/matte/tridymite equilibria in the Cu–Fe–O–S–Si system in controlled atmospheres: Experimental results at T = 1473 K [1250 °C] and P(SO2) = 0.25 atm. Metall. Mater. Trans. B. (accepted) 31. The lead and copper consortia database developed/optimized by PYROSEARCH, University of Queensland (2017) 32. Fallah-Mehrjardi A et al (2016) Phase chemistry study of the interactions between slag and refractory in coppermaking processes. In: Advances in molten slags, fluxes, and salts: proceedings of the 10th international conference on molten slags, fluxes and salts 2016. Springer 33. Lee WE, Zhang S (1999) Melt corrosion of oxide and oxide-carbon refractories. Int Mater Rev 44(3):77–104

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34. Fallah-Mehrjardi A et al (2016) Investigation of chemical interactions between slag and refractories in copper production processes. In: Copper international, Kobe, Japan 35. Fallah-Mehrjardi A, Hayes PC, Jak E (2014) Understanding slag freeze linings. JOM 66 (9):1654–1663 36. Fallah-Mehrjardi A et al (2013) Investigation of the freeze-lining formed in an industrial copper converting calcium-ferrite slag. Metall Mater Trans B 37. World economic forum (2017) Digital transformation initiative. http://reports.weforum.org/ digital-transformation/mining-and-metals-digital-transformation-and-the-industrys-new-normal/. Accessed 29 Mar 2018 38. McKinsey global institute (2016) The age of analytics: Competing in a data-driven world. https://www.mckinsey.com/business-functions/mckinsey-analytics/our-insights/the-age-ofanalytics-competing-in-a-data-driven-world 39. Gartner IT glossary. https://www.gartner.com/it-glossary/digitalization 40. https://deepmind.com/blog/deepmind-ai-reduces-google-data-centre-cooling-bill41. https://www.manufacturing.net/news/2018/01/q-inside-smart-steel-mill 42. Alvear Flores GRF (2014) Developments in copper smelting and refining a discussion of changes in copper smelting and refining industry in the last 40 years and some ideas looking into the future. ICSG, Lisbon 43. Copper 2013, Santiago, November 2013, Digital Procedures. Edited by IIMCH, Santiago, Chile 44. Yan J (2016) Recent operation of the oxygen bottom-blowing copper smelting and continuous copper converting technologies, PY16-2, Copper 2016, November 2016, Kobe, Japan 45. Cui Z, Wang Z, Wang H, Wei C (2016) Two-step copper smelting process at Dongying Fangyuan, PY16-4, Copper 2016, November 2016, Kobe;, Japan 46. Bing L (2016) Development of oxygen bottom-blowing copper smelting &converting technology, PY 21-3, Copper 2016, November 2016, Kobe, Japan 47. Michel S, Max S, Sumit M, Grit W (2017) Value-based production planning in non-ferrous metal industry—application in the copper industry. (Forthcoming)

The Changing World of Metallurgical Education Peter C. Hayes

Abstract The world continues to change and with it the supply of minerals and metals, the location of centres of production of primary metal and the increasing levels of metals and materials recycling. New technologies are being developed to meet the ongoing search by industry for lower costs, cleaner production and new markets. To keep abreast with these changes, and to utilise fully, the potential benefits of these technical advances, the industry will need a professional workforce having different knowledge, skills and professional attributes than in the previous millennium. What are these skills and attributes? How to best attract and develop the metallurgists of the future, and provide for the ongoing educational and research needs of the industry? Keyword Metallurgical education

Introduction The world needs metallurgists! The continued supply of the elements, most of which are metals, is critical for the sustainability of our technologically-based society. We need an industry that can deliver these resources, and we need educated people to design and efficiently operate the many, varied and complex production processes that make up the industry. So, what do we expect of the metallurgists of the future? What knowledge, skills and attributes should our future metallurgy graduates have? How can our educational systems deliver these outcomes? I have used the word “Metallurgist” to this point because some in the profession have Applied Science qualifications; most metallurgist now graduate as Metallurgical Engineers. The following discussion is focussed on metallurgical engineering. P. C. Hayes (✉) Metallurgical Engineering, School of Chemical Engineering, The University of Queensland, Brisbane, QLD Q4072, Australia e-mail: [email protected] © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_4

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Looking back, in the late 19th century Schools of Mines were established close to major mining operations to specifically educate the workforce on the technologies used in these operations; most of these mines and associated Schools have closed, those Schools that have survived have significantly changed focus and diversified into other fields of education. Minerals processing is by necessity undertaken at the mine site but practices have changed, metallurgists are now recruited from a broader geographical areas nationally and internationally. By the early 20th century, with the rapid increase in fundamental knowledge and understanding of physical systems, and the expansion of mass manufacturing technologies, metallurgical engineering education in Europe and North America covered a wide range of topics from mineral processing to the thermal treatment and manufacturing of steel. Primary metal production has peaked in these countries and Europe has embraced recycling and secondary metal production in a resurgence of activities to reshape their economies. In China, undergraduate education at individual Universities until recent years was based on the industry focus, e.g. Mineral Processing, Iron and Steel, Non-Ferrous Metallurgy; the offerings have broadened so these specialisations now only form a small part of individual university profiles. There has been a significant growth in the number of students and the proportion of the population undertaking tertiary education in most countries around the world, but these trends have not necessarily been reflected uniformly across the globe in metallurgy science or engineering. Looking to the future, in 2000 [1] predicted that the coming years will see, (i) A proliferation of information, (ii) Multi-disciplinary technological development, (iii) Globalised markets, (iv) Endangered environment, (v) Emerging social responsibility, and, (vi) Rapid change. There was another factor, (vii) Participatory corporate structures, while there is no clear evidence of that the latter has been achieved, however, the rest of the points seem to be well made and remain valid today [2]. We might add, (viii) the Digital revolution that is changing all aspects of technology and everyday life in a way and extent not previously envisaged. What educational programs do we need to develop for future metallurgists for the roles they may play in these scenarios? These are questions we need to ask before we design educational programs for the next generation and develop strategies to achieve these outcomes.

What? What will be the principal careers, career paths and roles for metallurgists in the future? As pointed out by [1], “the system of education is closely woven into the fabric of the society in which it operates”. For this reason the answers given to these questions will differ depending on when, where and who you ask. First, let us establish which engineers we are discussing. The primary qualification for engineers today is the baccalaureate (Bachelor’s) degree; this lays the

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foundation of technical skills used by engineers and recognition by professional societies. These technical skills are used principally in the early years following graduation as junior and plant engineers. The proportion of time spent in organisational rather than technical roles increases with exposure to plant practice and variety of operations. These management and economic roles are generally developed over longer timeframes (see Fig. 1). Starting at the first degree, most students study full time but are required to obtain some experience (usually in the long break between semesters) for a minimum time under the supervision of a professional engineer. Some undergraduate programs are structured to provide work experience through parallel cadetships or through COOP programs to provide opportunities for students to more fully develop graduate attributes and become more familiar with engineering practice. There are at present a limited number of “on line” rather than residential certificate and associate degree programs on offer but in general further study is required before full engineering accreditation is given to those following this path. The overall aim of most ME degrees currently offered is to provide advanced skills in a specific branch of engineering. These degrees are to prepare for technical roles in industry in design and operations. Single major Metallurgical engineering programs generally provide sufficient knowledge and skills levels for students to join the industry directly after graduation. Graduate training programs within companies generally provide a range of experiences in different parts of the operations to provide recent recruits with a broader appreciation of the business, and relationships between and interdependence of the different parts of the operation. It is desirable that new graduates receive mentoring, advice, guidance and instruction from Professional engineers within the company. Updates on industry

Engineering management

Management /economic

MBA/ CPD Technical support and sales

Company/plant operations

Professional engineering mentors Graduate programs

Cadetships COOP, Part-time studies Technical Vacation experience ME BE/BSc

CPD

MPhil PhD

Consultancies, Engineering design companies R&D, Process development Basic and Applied Research

BE (Met) Time

Fig. 1 Conventional career and learning pathways taken by metallurgical engineers over time

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trends and practice are usually provided in the form of continuing professional development (CPD); the extent to which this is undertaken depends critically on the company and management practice. Metallurgists seeking a career in industry research and development (R&D), consultancies and applied research in industry or at tertiary Institutions usually undertake a research higher degree (MPhil, PhD) to develop high-level specialist knowledge, and advanced technical and critical thinking skills. It is clear from these pathways that learning does not stop at the Uni gate. Again quoting from [1] “The education that succeeds will be the one that facilitates lifelong learning equipping students with the skills they will need to adapt to change”. A challenge for education providers and employers of the future is to integrate the career needs to options for educational advancement consistent with life-long learning models. This investment brings advantage to industry through establishing and retaining the corporate memory, providing career pathway and enhancing rates of retention of well-motivated employees.

How? Turning to the question of how can engineers develop the knowledge, skills, attributes and values for these roles? Knowledge As pointed out in a previous review [1, 2], over recent decades the “Knowledge base has grown so much that it is impossible for a single engineering curriculum to cover all and for graduate engineers to learn everything they before they graduate.” In addition, “engineers of tomorrow will work in a wide range of process operations throughout their careers”. How then to design engineering curricula that provide the educational framework and learning pathways? There are no right answers but many options—the solutions will vary depending on the context, industry and societal needs. • Continue to offer single-major specialised programs e.g. this is appropriate if there is a strong and sustained demand for a particular specialisation, e.g. Metallurgical Engineering. Primary metal production starts at the mine site and the International Mineral Processing Council (IMPC) has produced a road map recommending the suite of courses and skills that are desirable for degrees in Mineral Processing [3]. In some countries, particularly South America, Russia, parts of Asia, the standard model is 5–6 year BE, giving genuine breadth and depth to the programs. • Combine core engineering skills with other closely related degree specialisations in the form of dual major or dual degree programs. A range of options for these is illustrated in Fig. 2 showing potential relationships with related disciplines.

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An important feature of this option from a metallurgical engineering perspective is maintain the process metallurgy focus to the degree so that metallurgy is offered as a genuine major; to sustain this model strong internal support and collaboration within particular educational institutions is required. If metallurgy becomes a minor with limited breadth and depth, or is only available in introductory courses, the experience to date is that the program will struggle to survive in the long term in the University system; relying on the efforts of one or two academics is unstable and unsustainable. The aim should be to retain a critical mass; an academic complement with sufficient specialist skills and expertise to be able to genuinely offer a major program; this latter point is recognised in professional engineering criteria (see discussion below). • Produce generalist scientist/engineers BSc/BE (3 or 4 years) having core science and engineering fundamentals but with abilities to interact with, and be informed by, other disciplines. In this model, metallurgy would be only offered at the introductory level as part of a suite of electives. This provides preparation for subsequent coursework masters programs in specialist areas including Metallurgical Engineering if the student wishes to subsequently pursue further studies in depth. The danger here is that the ME program becomes effectively a re-labelled undergraduate program rather than providing genuine Masters level studies, creating confusion about the real level of competence attained. Some observations on the impact of the Bologna Model on mineral processing education are provided by [4].

Earth Sciences

Chemistry Physics Material Science

Mineral processing Mining Engineering

Metallurgy

Materials Engineering Mechanical Engineering

Engineering Management Economics

Chemical Engineering

Manufacturing Engineering

Fig. 2 Combinations of complementary and closely related fields of engineering knowledge and skills with metallurgy at the core

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Skills/Attributes To be able to design appropriate curricula to meet these needs it is useful to understand what has changed in engineering curricula over time—perhaps even since current practicing engineers graduated themselves. The focus of early engineering education models was on hands-on, practice-based curricula, focussed on the operation of technologies current at the time. This approach promotes information transfer and learning about how individual technologies work but does not provide adequate preparation for dealing with change. The significant changes to technology in the 20th century have resulted in commensurate changes in approaches to engineering education. These points and the thinking behind these changes are clearly explained by [5], who identified “five major shifts in engineering education that have occurred during the past 100 years: (1) from hands-on and practical emphasis to engineering science and analytical emphasis; (2) to outcomes-based education and accreditation; (3) to emphasizing engineering design; (4) to applying education, learning, and social behavioural sciences research; (5) to integrating information, computational, and communications technology in education.” Significant efforts have been made to define the skills and graduate competencies required by engineers. Typically the accreditation of engineers is overseen by Professional societies, e.g. American (ABET) [6], Engineers Australia (EA) [7], Institution of Chemical Engineers (IChemE) [8], China Engineering Education Accreditation Association (CEEAA) [9]. The criteria used in accreditation have changed overtime from rigid and prescriptive approaches, requiring originally definition of when and how learning takes place, to an “outcomes-based” accreditation approach—“using an outcomesoriented graduate capabilities standard against which the program is considered for accreditation; it does not specify the means by which these standards are met, giving the education provider freedom to design and execute programs” [7]. Similarly, the IChemE’s accreditation decisions are currently based an evidence-based assessment of the learning outcomes delivered by the degree programme and the levels at which these are achieved. In the USA ABET Criteria 2000 are the basis for accreditation. “As part of this assessment the program must have documented student outcomes that prepare graduates to attain the program educational objectives. The student outcomes are outcomes (a) through (k) plus any additional outcomes that may be articulated by the program.

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(a) an ability to apply knowledge of mathematics, science, and engineering, (b) an ability to design and conduct experiments, as well as to analyze and interpret data, (c) an ability to design a system, component, or process to meet desired needs within realistic constraints such as economic, environmental, social, political, ethical, health and safety, manufacturability, and sustainability, (d) an ability to function on multidisciplinary teams, (e) an ability to identify, formulate, and solve engineering problems, (f) an understanding of professional and ethical responsibility, (g) an ability to communicate effectively, (h) the broad education necessary to understand the impact of engineering solutions in a global, economic, environmental, and societal context, (i) a recognition of the need for, and an ability to engage in life-long learning, (j) a knowledge of contemporary issues, (k) an ability to use the techniques, skills, and modern engineering tools necessary for engineering practice.” Other attributes include, problem solving, critical and creative thinking, interpersonal and teamwork communication skills, integrative and systems thinking and change management skills [1]. As metallurgical engineers we understand and appreciate that the technological processes we operate are complex, and potentially can cause harm if they are not designed and controlled appropriately. Given the importance of these issues, understanding how to think about safety; environmental, personal and economic risk, impact and risk management, professional responsibility and ethics, and how to decide on appropriate actions and outcomes associated with the issues, I personally believe should be part of graduate engineering education in the 21st century. The future will see greater quantitative description of processes using the rapidly increasing computing capacity and speed. This opens the possibility of further improvement of process efficiencies through better control and optimisation. All process plants are complex networks of process streams. To be able to understand and describe in quantitative terms the interrelationships between processes, and the dynamic systems in which they operate, points to the need to integrate specialist metallurgy skills and expertise with the “systems thinking” approach developed in chemical engineering. The systems approach and with it the ability to analyse and model process flow sheets adds an important new dimension to metallurgy programs. A dual BE major in Chemical and Metallurgical engineering provides the core engineering science, and process engineering knowledge and skills associated with conventional chemical and metallurgical engineering programs; this is the approach that has been adopted at The University of Queensland [10]. Post-graduation education and enhancing the competencies of practising professionals is a more difficult area to summarise and one that remains uncertain since there are a wide variety of career pathways.

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Despite the availability of jobs in industry, there have been shortages in personnel with metallurgical technical skills over the past several decades. The shortages have been addressed by the industry in various ways [11], some more successful than others. In addition to encouraging increased BE enrolments options for further education include; Coursework Masters general (certification); Coursework Masters focussed on particular industry sectors (certification); Conversion courses for engineers without metallurgy background (certification); Short course formats (attendance); Conference (certification) [7]. These approaches address to some extent the short-term needs of industry and plant operations; but they do not necessarily go to the heart of developing expert skills and professional competency. In particular, whilst short course format enables the transfer of information it does not provide the opportunity for practice and feedback, which are essential elements for “deep learning”, and long term retention of and competence in the subject matter. This limitation should be recognised by industry- short term fixes can be deceptive and are not always the best solution. The very significant decline in the number of experienced process engineers at operating sites is a serious constraint on the extent to which young professionals can develop under the guidance of experienced engineers, even though some of the larger corporations have well-structured generic graduate programs. The term professional engineer describes a person holding an engineering qualification from a university degree course accredited by an engineering profession e.g. Engineers Australia, and who has undergone a period of formation in the workplace. However many university graduate engineers do not join any professional organisation, and many do not even go on to practice engineering in a professional or technical sense. Some countries and jurisdictions require that professionals must be registered or chartered, however, this is by no means a universal requirement. Since the activities are many and varied it should be the responsibility of each engineer assess the professional knowledge, competencies and the relevant training required to undertake these tasks. The Warren Centre report [12] identified eight essential elements of performance when acting in a professional engineering capacity. “The Professional Engineer should • Develop a clear understanding of the Relevant Parties to and Other Stakeholders in the Engineering Task and the relationships between them. • Consult and agree with the Responsible Person the objectives and extent of the Engineering Task. • Assess and apply the competencies and resources appropriate to the Engineering Task. • Identify and respond to relevant statutory requirements and public interest issues. • Develop and operate within a Hazard and Risk Framework appropriate to the Engineering Task. • Seek to use engineering innovation to enhance the outcomes of the Engineering Task.

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• Apply appropriate engineering task management protocols and related standards in carrying out and accomplishing the Engineering Task. • Ensure that any contract or other such evidence of agreement governing or relevant to the Engineering Task is consistent with the provisions of this PPIR Protocol.” Using these guidelines may help individuals identify the professional engineering attributes that are necessary for their practice.

Where? Where can you study metallurgical engineering? The short answer is, in relatively few Universities in industrialised countries, and the coverage is not uniform across the globe. There was a significant decline in the number of metallurgical engineering programs in Europe and North America in the late 20th century in parallel with changes to the industry profile; as primary metal production in those countries decreased, and the investment and teaching resources were moved to materials science. Iron and steel production expanded in Asia progressively through Japan, Taiwan, South Korea, China. After an initial decline there has been a recent renaissance in Europe with the impetus provided by the changes in industry profile, the need for resource security and sustained supply of new elements/materials, recycling, reprocessing and the implementation of new process technologies to address environmental issues [13]. In Australia, despite the major contributions to the economy by the minerals industry there are low domestic student enrolments in metallurgy. Recent decades have seen a decline in metal processing and refining, and an increased reliance on mineral concentrate exports. Recruitment into undergraduate metallurgy programs in Australia has been shown to be directly related to metal price, and subject to the wild fluctuations of the business cycle. With the expansion of participation in tertiary education and increases in standards of living across the globe the numbers of international students undertaking engineering studies in North America, Europe and Australia have dramatically increased. A recent National Foundation for American Policy survey [14, 15] found that in the USA “International students make up the large majority of full-time students in many graduate science- and engineering-related programs, and their numbers have been rising much faster than the number of domestic students”. International students were found to make up 50% of enrolments at undergraduate engineering programs, and even greater proportion at coursework Masters and PhD levels. These trends are also increasingly reflected in metallurgical engineering programs. Whilst there have been declines in metallurgy programs in industrialised counties elsewhere, where mineral exports have increased, the establishment of new

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metallurgy programs has taken place; governments have actively promoted employment of local workforces rather than expatriates workers, which was the predominant model for many years. In China, there have been major investments in education infrastructure and capability in all branches of science and engineering in parallel with the expansion across all aspects of this economy [16]. It is one thing to design appropriate programs, it is another to persuade students that metallurgical engineering is the career for them. In an age where choice of degree programs on offer has blossomed, it is difficult to attract students into the discipline; key factors influencing enrolments are Careers, Pay, Relevance, Flexibility. Just as individual degree programs are tailored to meet academic and technical requirements finding and establishing a profile that is attractive to prospective students is a major issue. Teaching academics work hard at attracting students into their programs. Strange as it may seem publicity for metallurgy programs are not always actively supported by University administrations, who see the advantages of economies of scale associated with large class sizes in other major disciplines. In Australia, the exaggerated boom/bust commodity price cycles associated with primary metal production constantly erode student and parent confidence in the industry; over the past six decades undergraduate enrolments in mining and metallurgical engineering are a clear function of metal price and graduations have been out of cycle in almost all cases. This is not a good way to operate a process and not an efficient method of developing a high quality, skilled professional workforce. It is my view, that the industry needs to demonstrate its commitment and ongoing support over the longer term, and provide sustained and attractive career opportunities if it is to attract the educated workforce it needs for its future.

Summary Just has the profile of the metallurgical industry has changed over the years so too has metallurgical education in terms of the places that provide these learning opportunities, the curricula offered, the approaches to learning and connections to industry. Important questions for industry and education providers are • What specialist knowledge and skills, and attributes are we looking for in the metallurgists of the future? • What level of qualification are we seeking? Graduate BE, or advanced standing Masters, PhD, and /or other? • How can industry help to provide and sustain these educational opportunities? • How to best attract these potential employees into the profession?

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References 1. Rugarcia A, Felder RM, Woods DR, Stice JE (2000) The future of engineering education: Part 1. A vision for a new century. Chem Eng Education 34(1):16–25 2. Wankat PC, Bullard LG (2016) The future of engineering education revisited. Chem Eng Educ 50:19–29 3. Drinkwater D (2018) International mineral processing congress, IMPC mineral processing education roadmap. http://impc-council.com/wp-content/uploads/2018/02/IMPC_Mineral_ Processing_Education_ROADMAP_opt.pdf 4. Batterham R (2012) The impact of the Bologna model on mineral processing education: good, bad, or indifferent. In: Minerals industry: education and training. Proceedings of the 16th international mineral processing congress (IMPC), Institute of Mineral Engineers, Indian, paper 1104, pp 390–395 5. Froyd JE, Wankat PC, Smith KA (2012) Five major shifts in 100 years of engineering education. Proc IEEE 100:1344–1360 6. Accreditation board for engineering and technology (ABET), engineering accreditation commission. www.abet.org/ 7. Engineers Australia (EA). www.engineersaustralia.org.au/ 8. Institution of Chemical Engineers (IChem E). www.icheme.org/ 9. CEEAA, China engineering education accreditation association. http://english.ceeaa.org.cn 10. Hannah B, Hayes PC (2014) The challenges for professional metallurgical education. In: Celebrating the megascale: extraction and processing symposium on pyrometallurgy in Honor of Robertson DGC—Mackey PJ, Grimsey EJ, Jones RT, Brooks GA (eds) TMS 143rd Annual Meeting, San Diego, TMS, pp 473–480 11. Drinkwater D (2012) Special symposium on human resource development. In: 16th International mineral processing congress (IMPC). Indian Institute of Mineral Engineers, New Delhi, India 12. Warren centre (2009) Professional performance, innovation and risk in Australian engineering practice. The Warren Centre for Advanced Engineering, Sydney 13. Metallurgy made in and for Europe (2014) The perspective of producers and end-users roadmap. ISBN 978-92-79-43310-8, European Union. https://doi.org/10.2777/11914 14. Young C (2016) Montana tech: perspectives of a small specialty school, IMPC Roadmap 15. National foundation for American policy (2017) http://nfap.com/wp-content/uploads/2017/10/ The-Importance-of-International-Students.NFAP-Policy-Brief.October-20171.pdf 16. Hu Y (2016) Status and prospects of mineral processing postgraduate education in China, IMPC Roadmap

Part II

7th International Symposium on Advances in Sulfide Smelting

Sulfide Smelting Development in Japan During the Past Half Century Takahiko Okura and Hiromichi Takebe

Abstract Japanese non-ferrous industry introduced large-scale smelting plants along seashores for overseas concentrates with increasing demand of metals around 1970s. Although serious environmental pollution became obvious, with economic growth, the industry got rid of pollution by installing new processes and improving operation technologies. At present, over 99.8% of sulfur input to smelters is fixed as stable compounds. Through those decades Japan had steep rises of oil prices, sudden change of exchange rate, and inadequate treating charges, we were faced to consider the closure of smelters. The industry has survived by increasing productivity, saving energy and reducing manpower. Furthermore the industry made great effort to recycle valued metals from scraps and wastes for the resources-recycling society. Academic research also contributes to support these individual technologies. Thus the industry has fostered world-acclaimed technologies in terms of efficiency and energy conservation. This paper presents technology development and environmentally-benign sulfide smelting processes.





Keywords Japanese non-ferrous industry Cu smelting Zn smelting Pb smelting Ni smelting Environment Sulfur Recycling Thermodynamics Kinetics











Introduction Japanese non-ferrous industry operated blast furnaces, reverberatory furnaces and so forth with self-developed technologies using concentrates mainly from domestic mines after the World War II. As the increase of demand along with economic development around 1970, the industry introduced large-scale smelting processes T. Okura (✉) Ehime University, Tokura 4-23-16, Kokubunji, Tokyo 185-0003, Japan e-mail: [email protected] H. Takebe Ehime University, Bunkyo-cho 3, Matsuyama, Ehime 790-8577, Japan © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_5

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along seashores for overseas concentrates. Various troubles faced at the initial period had been solved through many technological improvements. Although serious environmental pollution became obvious, together with economic growth, the industry got rid of pollution by installing new processes and improving operation technologies. At present, over 99.8% of sulfur input to Cu smelters is fixed as stable compounds such as sulfuric acid and gypsum. It means that all of the smelters in Japan are operated environmentally friendly. Through those decades Japan had steep rises of oil prices, sudden change of exchange rate, and inadequate T/C & R/C (treating/refining charges) rates, and the industry were faced to consider the closure of smelters. The industry has survived by increasing productivity, saving energy and economizing on manpower. Furthermore the industry made great effort to recycle valued metals from scraps and wastes in order to contribute the resources-recycling society. Thus Japanese non-ferrous industry has fostered world-acclaimed technologies in terms of efficiency and energy conservation performances. Increase of flash smelting productivity, the Mitsubishi Continuous Copper Smelting Process (the Mitsubishi process), the hematite process for zinc residues, and SMM’s Ni hydrometallurgical process are worthy of special processing technologies. Along with the industry effort, academic research also contributes to support these individual technologies. The examples of the contributions are the Yazawa potential diagram, and metallurgical studies on the impurity distribution between matte and slag and the ferrite slag. Some of the contents are described in “The history and tradition of Japanese non-ferrous smelting technology [1]”.

Historical Conspectus of Japanese Non-ferrous Industries Annual production of base metals were some 30 kt of copper, 10 kt of zinc and 5 kt of lead respectively just after the World War II. They were gradually increasing with economic recovery and reached, till 1960, 250 kt of copper, 180 kt of zinc, 180 kt of lead and 5 kt of nickel mainly using domestic concentrates as shown Fig. 1. Then Japan had the steep economic growth. It forced the industry to enhance the production rate, and introduced the new smelting processes. They are listed in Table 1. As a characteristic, some of them were joint venture among the industry. Teething problems were solved through in-house metallurgist and engineers together with academic researches. After stable operations were attained, Japan had severe change of business circumstance such as the hikes of oil price, higher appreciation of Japanese yen to US dollars, lower T/C & R/C and so on. The industry has changed their business model properly and survived. Meanwhile, the air and water pollution became severe problematic during 1960s and 1970s. Regarding sulfur oxide gas emission, following stricter governmental regulation, complete fixation were pursued. The fixation ratio is now more than 99.8% in Cu smelters. By the way, early 20th century, smelting gas was dispersed to the atmosphere through a high stack to decrease the damage to the ground around

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Fig. 1 Production of copper, zinc, lead and nickel, 1960–2017

Table 1 New smelting processes around 1970s [DMM] DOWA Metals & Mining Co., Ltd. [FCL] Furukawa Co., Ltd. [JXM] JX Nippon Mining & Metals Co., Ltd. [MMC] Mitsubishi Materials Corp. [MMM] Mitsui Mining & Metals Co. Ltd. [SMM] Sumitomo Metal Mining Co., Ltd Smelter Cu

Ashio Kosaka Onahama Saganoseki Hitachi Saganoseki Toyo Tamano Naoshima Zn Akita Hikoshima Zn– Harima Pb Hachinohe Ni Niihama (unit; kilotonnes/year)

Process

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1956 1967 1965 1970 1972 1973 1971 1972 1974 1972 1970 1966 1969 1989

12 36 72 120 84 120 100 84 48 78 55 36–18 54–24 25

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FCL DMM MMC etc. JXM JXM JXM SMM MMM MMC DMM etc. MMM SMM MMM etc. SMM

smelters. They are still working. however, one of the symbols in old days that was No. 1 Stack of Saganoseki Smelter and Refinery (167.6 m and the highest in the world at the time), was demolished ending its mission safely and completely [2]. Sulfuric acid plants were also enlarged according to smelting capacity, whose main

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process is DC/DA (double contacts and double absorptions), and some of them have one of the largest capacities in the world. Anodic protection cooling process for irrigation cooler using sea water is one of specified technologies in acid plants.

Copper Smelting Technologies The trend of copper production would be divided into 3 phases seeing Fig. 2, namely developing and stabilization phase, enhancing of production, and improving technologies & establishing new business models. After the copper industry had developed its own technologies such as Momota process at blast furnace and oxygen-smelting in PS converter followed by converter slag flotation [3] which has developed as the Teniente smelting process, the industry needed new bigger production process to satisfy the demand, and commenced the operation of Outokumpu flash smelting furnaces and the Mitsubishi Process as shown in Table 1. Currently the industry exports some 30% of the production mainly to China. PHASE—1: Developing and stabilization of the introduced processes between 1960s and 1980s FCL had introduced the first flash smelting process outside Finland in the world. Their efforts and trials very contributed to the development of the process. Following Ashio Smelter, 6 flash furnaces began operation. JXM used hot blast of 1,000 °C with small amount of oxygen enrichment, and MMM constructed the flash furnace with electrodes inside its settler, which developed CO/CO2 concentration controlling model for decreasing the Cu content in the slag. MMC started the Mitsubishi Process after 17 years great effort and the invention of the ferrite slag

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composition (CaO–Fe2O3–Cu2O) [4] that has been very conventional at flash converting process nowadays. The industry increased the production rate and exported the technology to the world. During this era, it began to use tonnage-oxygen to reduce the energy cost due to higher oil price, and had to actively seek the resources from overseas when the ratio of domestic concentrates decreased below 30% in 1960s. As a custom smelter, the industry had a role of main player in the world concentrate market, and developed sophisticated blending technologies. PHASE—2: Enhancing production during 1980s–2005s SMM invented its own single burner for flash smelting followed excellent study on the behavior of concentrates and reaction gas in 1996 [5], and finally enhanced the production rate to 450 kt/year in 2007 [6]. JXM integrated two flash furnaces into one furnace [7], and gradually increased the feed rate finally to 220 t/h and attained the full capacity of 450 kt/year in 1998 [8]. MMC succeeded the enlargement of the Mitsubishi Process in 1991 followed by delicate reaction and process analysis. Its current capacity is 300 kt/year. PHASE—3: Technology improvement and business model developing for survival after 2000s Onahama smelter developed new fuel resource from waste tires and the stable combustion technique for automobile shredder dust, and finally it set up S-furnace of the Mitsubishi Process to increase the production [9]. Kosaka quit flash smelting for imported concentrates, and changed the raw materials to valued metal containing scrap together with entire process change, which is composed of TSL— blister Leaching—EW [10]. It also constructed landfill site in the trace of old copper mine with excellent waste water treating process. In the area of environmental solutions, smelting processes have closed circuit of granulation water for its slag, and de-sulfurization process for low SOx gas was invented. Converting: Campaign life is the key factor for efficient operation, which has been attained through precise and stable operation using controlling/simulation models and so on. Saganoseki reduced the number of PS–converters from 6 to 4 in 2005 and set the new charging facility for anode scrap, Toyo meanwhile constructed one more converter and attained 3 hot—2 blow scheme [11]. Their time schedule in converter aisle is very efficient as written in references. During this phase, waste gas boilers were introduced to recover energy and for tight gas handling. Electrolysis: In Phase-1, Onahama commenced refining in the big cell with continuously cast anode, and Tamano initiated PR (periodic reverse)—electrolysis. In phase—3 came, ISA process was introduced at Hitachi Refinery of JXM in 2002 followed by Toyo, Saganoseki, and Tamano [12]. SMM operated it at 350 A/m2 after explicit study on passivation [13]. According to increase of scrap metals, Sb and Bi loads exceeded the capacity of existing facilities, then the chelating resin absorption process was introduced to many refineries and being reinforced. Anode Slime treatment: Conventional process has longer recovering period of gold. JXM developed the hydrometallurgical process in 1997, and SMM and MMC

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commenced their operations in 2004 respectively due to some reasons such as decrease of interest, decrepit facilities, and environmental problems [14]. Recycling: Copper is an excellent student in recycling world due to its characteristics. Unit consumption of copper per population in Japan has matured and it disperses into various wastes as a small amount. Contributing to the resource-recycling society and survival of the industry, it is treating scraps and valued metal—containing wastes as much as possible. They have added new facilities such as incineration—melting kilns, new sampling facilities and so on.

Zinc and Lead Smelting Technologies The present business scheme of Zinc industry has been established in 1970s, however, the total production is decreasing gradually due to the decline of domestic demand and insufficient T/C & R/C terms as shown in Fig. 3. During these tough periods, the industry devoted themselves for technology development and new returns. At the beginning of the era, Akita Zinc introduced the hematite process for residue treatment, and Hikoshima Smelter began electrolytic Zn winning process at different Dk (current density) between daytime and nighttime in order to decrease the power cost because it is cheaper in nighttime than in daytime. This operation scheme has become standard in Japan. As other technologies worthy of special mention, electro-winning at higher Dk, new composition of Pb-anode, and the technology of very low Pb-containing cathode [15] should be appreciated. Furthermore, roasting technologies for high-Si concentrates and very fine concentrates are established recently [16]. 70

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At ISP of Hachinohe, some excellent technologies such as the latent heat recovery of gaseous zinc, hot-briquetting of zinc oxide, fuming of slag for ZnO recovery, and enlargement of the furnace area to the world largest one were developed. Toho Zinc is producing ZnO at the electrothermic distillation process in Onahama Smelter [17]. Akita Zinc Recycling is operating the plant of 30 kt/year, whose process is composed of HCl leaching of wastes, solvent extraction, stripping by return electrolyte from main tankhouse and EW there. As shown in Fig. 3, the lead industry has shifted to waste Pb-battery recycling following the decrease of domestic concentrates, contributing to prevent the hazardous element release to the environment. Each smelter is developing own technologies such as high Ag-anode electrolysis [18], and introduction of an incineration-melting kiln for waste print circuit boards.

Nickel Smelting Technologies SMM has a long history of nickel smelting technologies since 1930s. SMM is continually enhancing the productivity of Ni smelting as shown Fig. 1. In 1970 matte-electrolysis technology was established. The worthy technology is so called MCLE (Matte Chlorine Leach Electro-winning) process that started in 1989, whose process flowsheet is composed of chlorine leaching, de-copperization in chloride bath, solvent extraction, and EW of Ni [19]. After SMM commenced the operation of the HPAL process in Philippines, they had to improve the process for the low-leachable mixed sulfide from the plant [20]. Meanwhile, Fe–Ni alloy are produced from Ni oxide ore in Japan, too.

Academic Contribution to the Industry Potential diagrams, invented by Professor Yazawa [21], is the mile-stone for the metallurgists in the world, that explains the Cu smelting process thermodynamically, followed by impurities distribution between matte and slag including Cu loss in silicate and ferrite slag. In the area of kinetics, CFD (computational fluid dynamics) models for reaction and particle behavior in the flash smelting are making them clear. Recently some governmental institutes have many projects to study on treating high as containing ore and/or concentrate, and recycling valued metals, in which the industry respects their support. Collaboration among the industry, the academic institute and the government is being executed accordingly to the epoch change. It is necessary for us to succeed and develop legend technology even though almost of conceptual theory and technology have been become clear.

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Conclusion Japanese non-ferrous industry experienced many float and sink in its history due to such as increase of domestic demands and worse market conditions. All of metallurgists are still devoting themselves to enhance technology. Recently the industry spreads its world to South-Eastern countries to survive together with technologies transfer. The current main objective is to smelt low grade and high impurity concentrates efficiently. It shall emerge in the quite near future. Acknowledgements The authors thank Professor Emeritus Masuko and Professor Emeritus Yamashita who edited reference [1] on their excellent initiative. Mining and Metals Association of Japan’s support for making the numerical history is also the great help for us. We finally and sincerely thank all of the metallurgists in the industry and the academic institute for their legend and continuing technical developments.

References 1. Masuko N, Yamashita S et al (2005) The history and tradition of Japanese non-ferrous smelting technology, research report (in Japanese). Jpn Soc Promot Sci 2. Nishi T, Kawano Y, Arakane T (2016) Demolition of the no 1 stack at Saganoseki smelter and refinery, Copper 2016, MMIJ, PY17-4 3. Tsurumoto T (1961) Copper smelting in the converter. J Metals 13(11) 4. Takesue T (1981) Copper smelting at Tamano smelter. J MMIJ 97:637–642 5. Kemori N, Denholm WT, Kurokawa H (1989) Reaction mechanism in a copper flash smelting furnace. Metall Trans AIME 20B:327–336 6. Otaka S, Kobayashi J, Yamamoto K, Kawanaka K, Mori K, Sasaki Y, Hattori Y (2016) Flash smelting operation at the Toyo smelter and refinery, Copper 2016, MMIJ, PY2-4 7. Ishikawa M (1998) J MMIJ 114:447–454 8. Soma Y, Hong JH, Motomura T, Chida H (2016) Improvements to the Saganoseki flash smelting furnace operation, Copper 2016, MMIJ, PY2-1 9. Kiyotani K, Iida O, Tanaka F Copper 2010, GDMB, 1199–1210 10. Ogawa K (2016) Copper electrowinning and nickel recovery from black copper containing high levels of impurities from the smelting process of complex recycling materials, Copper 2016, MMIJ, EL7-1 11. Papers on PS-converting process, Copper 2016, MMIJ, PY7-1 & PY7-2 (2016) 12. Nagao S, Shikada K, Yamaguchi Y, Kanazawa M (2016) Recent operational improvements at Saganoseki electrorefinery, Copper 2016, MMIJ, EL1-4 13. Toki N, Akiyama T, Nagase N, Hattori Y (2016) High current density operation of conventional copper electrorefining, Copper 2016, MMIJ, EL1-2 14. Papers on hydrometallurgical slime treating process, Copper 2016, MMIJ, HY9-2 & HY9-4 (2016) 15. Nakamura H, Yamada K, Aichi T (2015) Development of the “Pb Less” electrolytic zinc technology. In: Proceedings of Pb–Zn, GDMB, 1077–1086 16. Yoshida T, Nishijima A, Ono T, Nishiyama F, Baba T (2015) Process modification to increase the amount of fine concentrates by wet feed type roaster in Hikoshima zinc smelter. In: Proceedings of Pb–Zn, GDMB, pp 593–600 17. Sugawara A (2015) Zinc and lead smelting at Hachinohe smelter. In: Proceedings of Pb–Zn, GDMB, pp 517–525

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18. Kitano T, Shima N, Hashimoto N (2015) Effect of various conditions in the lead electrolytic process. In: Proceedings of Pb–Zn, GDMB, pp 375–388 19. Ishikawa Y (1994) J MMIJ, 110, 913–918 20. Sato H, Sakamoto K, Ooishi T, Sato Y (2017) Increasing capacity of nickel production at Niihama nickel refinery, Nickel–Cobalt 2017, CIM (2017), Paper No. 9425 21. Yazawa A (1979) Thermodynamic evaluations of extractive metallurgical processes. Metall Trans AIME 10B:307–321

Review of Boliden Harjavalta Nickel Smelter Hannu Johto, Petri Latostenmaa, Esa Peuraniemi and Karri Osara

Abstract Boliden Harjavalta has operated a flash smelter for nickel sulphide concentrates in Finland since late 1959. The original nickel flash smelting process was modified in 1995 to a novel nickel matte smelting (DON) with semi-low iron where the batchwise operated converting became obsolete. Since that, the process has been applied to multiple chemically variable nickel sulphide concentrates with high environmental performance. In this paper, the status of the current smelter operations is described together with the recent modernizations of the nickel flash smelting furnace and the slag cleaning furnaces. Keywords Boliden



Nickel smelting



DON

Introduction Boliden is a metal refining company with operations located mainly in northern Europe with core competence in the fields of exploration, mining, smelting and metal recycling. Boliden Harjavalta, located in Finland, is a part of Boliden Group and operates nickel and copper smelters and sulphuric acid plants in Harjavalta, and copper refinery in Pori. Flash smelting of copper concentrates was commissioned in Harjavalta in 1949. A decade later, nickel flash smelting was commissioned, in 1959. Due to the development of furnace campaign life, there was no more need for initially two copper smelting furnaces and thus the other could be utilized for nickel concentrate smelting. The major drivers for the development of a new smelting process during 1940s and 1950s were energy efficiency and high sulphur recovery, which are also today important factors in smelting of sulphidic concentrates. Today, flash smelting technology represents around 50% of world’s annual primary nickel production from sulphidic raw materials [4]. At first, the flash smelting—Peirce-Smith converting route was applied at Harjavalta nickel smelter. Slag from the both flash furnace and the converters was H. Johto ⋅ P. Latostenmaa (✉) ⋅ E. Peuraniemi ⋅ K. Osara Boliden Harjavalta, Teollisuuskatu 1, 29800 Harjavalta, Finland e-mail: [email protected] © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_6

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treated in an electric slag cleaning furnace. In the 1980s, a development project for nickel flash smelting process without converters was initiated. The aim of the project was to create a process capable of: • reducing the operating and maintenance costs by replacing the Peirce-Smith converters • making the flash smelting viable for concentrates with high MgO, low copper and high PGMs • reducing emissions from the smelter • improving the working atmosphere In commercial scale the new process was commissioned in Harjavalta with expanded production in 1995; the process was named as Direct Outokumpu Nickel (DON) Smelting process [3]. The flow sheet of the process is shown in Fig. 1. Today, Harjavalta smelter is the only operational nickel smelter utilizing the DON process. Historical concentrate feed data during the nickel smelter’s operational time is shown in Fig. 2, indicating also the feed record year in 2016. For historical reasons, Harjavalta nickel smelter has performed toll-smelting for the near-by refinery located at the same industrial area. However, in mid 2015, Boliden changed the nickel business model and initiated its own concentrate supply and sale of matte to market.

Fig. 1 The flowsheet of Boliden Harjavalta nickel smelter

Fig. 2 Annually treated nickel concentrate at Harjavalta nickel smelter. 2016 was the production record year

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Raw Materials Boliden Harjavalta nickel smelter operates with complex Ni–Cu bulk concentrates with variating nickel and copper grades. The DON process is capable of processing a variety of concentrates as described also elsewhere [3]. High magnesia in the feed mixture strongly influences the flash smelting slag properties. Due to the higher oxidation of feed mixture’s iron, MgO can be more efficiently diluted in the slag of the DON process compared to the conventional nickel flash smelting process. Thus, the operational temperature of smelting vessel can be kept much lower than in the conventional flash furnace with the same feed mixture Fe/MgO-ratio. Since Boliden’s acquisition of Kevitsa Ni–Cu–PGM mine in northern Finland in 2016, the main part of the smelter feed has consisted of Boliden group’s own concentrates. Kevitsa Ni-concentrate with relatively high MgO content together with moderate Ni/Cu-ratio and PGMs contained is a well suitable addition to Harjavalta’s concentrate mixture to be treated in the DON flash furnace. Concentrates are shipped to Pori harbour where as the smelter is located only 60 km inland from the port. From the harbour, the concentrates are transported mainly by rail with tailor made covered wagons to Harjavalta concentrate storage.

The DON Process and the Latest Modernization of the Flash Furnace The blended concentrate mixture is dried using an oil fired rotary drum with capacity up to 60 t/h. The blend is made from feed bins by adjustable belt feeders together with the needed silica sand as a flux former. The dry charge is then conveyed by airlift to dry charge bin. During the major overhaul in 2017, an Outotec loss-in-weight feeder was installed to replace the old drag conveyer as a concentrate feeding system. With bone-dry and fine feed blend, the drag conveyer had difficulties to maintain constant feed rate causing problems in the furnace control. With more accurate feed rate control of the loss-in-weight system, a better control of the furnace performance was achieved. The oxygen utilization with the new feeding system has improved. The DON process oxidates concentrate feed in the flash furnace further in single step compared to conventional nickel flash smelting process. This results a matte with relatively low iron (4–6% Fe) and high in valuables content, around 70% Ni + Cu + Co. Highly oxygen enriched combustion air is used as a blast. A single concentrate burner equipped with an air slide was installed already in 2008. Matte is granulated directly from the furnace utilizing Uddeholm’s Granshot technology. Harjavalta nickel flash furnace was fully rebuilt in 2017. The new furnace design is shown in Fig. 3. The old furnace steel frame originating from 1991 was totally demolished. The settler bottom cooling is required to prevent low-in-copper matte from penetrating into the bricklayer. Forced air cooling is used; it was modified

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Fig. 3 Modernized nickel flash furnace

from uniform cooling to sectional cooling in order to improve control possibilities especially in case of bottom build-up formation. Heat transfer efficiency of all the three cooling sections can be adjusted to control the bottom cooling independently between the reaction shaft and the uptake shaft. Main target of the modernization was to extend the furnace campaign life from two years to four years. The slag-matte interface on settler walls required a partial re-bricking every two years. In order to reach the target, water cooling of the furnace was increased and enhanced to reach longer refractory lifetime. Altogether, some 300 tons of copper cooling elements were installed together with more than 6 km of stainless steel piping supplying the cooling water. Cooling of the reaction shaft was changed from the obsolete water film cooling to modern copper cooling element structure covering the whole reaction shaft in three layers. Based on the experience from Boliden Harjavalta’s copper flash furnace [2] and the CFD modelling during the engineering [1], the design of the uptake shaft was modernized from its original rounded rectangular shape to a round one. At the same time, the diameter of the uptake shaft was decreased and its copper cooling design was modified. One of the furnace weak points had been the lifetime of the lowest cooling elements in the transition zone between the settler roof and the uptake shaft; special emphasis was paid to this area to improve its durability.

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The overall duration of the furnace rebuild from feed-off with the old furnace till starting the feed again with the modernized furnace was 45 days. The planned strict time schedule was well followed. An additional challenge during the rebuild project was the copper smelter that was operating just next to the nickel flash furnace under modernization. During this maintenance work, one LTI accident was recorded. The magnitude of the rebuild work can be seen in Fig. 4. Since the rebuild, the new flash furnace feeding system has been working without major disturbances. The flash furnace start-up after modernization was performed better than expected and the full capacity was reached within couple of days. Lately, an excessive build-up formation at the boiler throat has been experienced influencing the production; full understanding of the root causes is not yet known while this paper is submitted. Currently, the flash furnace is operated with an iron-in-matte target of 6% as this have been found suitable from both the customer and the smelter recovery perspective. The target iron content is higher than the smelter traditionally has operated. A wide range of investigations together with Aalto University has been conducted over the years; the results are described elsewhere in this conference [5]. The studies deal with the matte/slag distribution behaviour of different metals in flash smelting conditions. The new knowledge on the process collected within these studies combined with on-site work has increased the ability to predict smelting behaviour of different concentrates and thus, the product quality. The high oxidation degree in the flash furnace causes losses of valuable metals into flash furnace slag. In Harjavalta, the slag is cleaned by coke reduction in an electric furnace fully rebuild and modernized in 2014 [2]. From the electric furnace, a sulphurized metallic alloy is tapped and water granulated using similar process than for the flash furnace matte; the alloy is sold separately to customers as EF matte. A new copper made slag launder from flash furnace to slag cleaning furnace with overlapping joints was also installed during the 2017 rebuild project. This aimed to improve operational safety by limiting the possibility of slag leakages from the launder joints. Since the installation, no melt leakage from the connections have been experienced.

Fig. 4 The nickel flash furnace modernization; new bottom steel structure installed and first new settler wall copper cooling elements on their locations on the left. Photo taken from the matte taphole end looking towards uptake shaft/boiler. On the right, the new reaction shaft copper cooling without brick lining under construction

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Environmental Considerations Boliden Harjavalta develops its environmental performance continuously. The main stack emissions points are monitored by on-line measurements and the state of the environment is followed by real time measurements at environmental stations in the surrounding community. Emissions are monitored and controlled in close cooperation with supervising authorities. Based on environmental performance our plant manages well in international comparison. The DON process is accepted on the BAT reference list of European Union as it represents the best available technique in reducing emissions and energy consumption of metal production. The smelting process is continuous and ladle transportation to batch wise operated converters is not needed, the few emission sources are easily covered and ventilated. This leads to well controlled fugitive emissions from the plant. The off-gas from the flash furnace is directed to sulphuric acid plants after heat recovery boiler and hot-ESP. A new, modern sulphuric acid plant will be fully commissioned in 2019 replacing the oldest operational sulphuric plant in Harjavalta. This replacement improves sulphur capture even further. The ventilation gases from the flash furnace matte tapping and the electric furnace alloy tapping are treated for SO2 fixation using sodium hydroxide based venturi scrubber. Figure 5 shows historical SO2 emission data as specific sulphur emissions (kg SO2/ton metal copper + nickel produced). The cleaned slag from the electric furnace is water granulated. Slag production is about 200 000 tons per year and the slag is classified as normal waste according to Finnish and European standards. Part of the slag is utilized for sandblasting abrasive and in building material industry but the major share of the slag is landfilled to the smelter area.

Fig. 5 Harjavalta’s sulphur dioxide emissions from nickel and copper smelters

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Conclusion Boliden is committed to long-term sustainable and environmental responsible development of its operations. Boliden Harjavalta has a long tradition in operating a nickel flash furnace. In 2017, a full modernization of the flash furnace was performed in a very short time of only 45 days. The modernized smelting furnace together with other improvements such as new acid plant investment in Harjavalta indicates the commitment Boliden has on sustainable nickel production also in the future.

References 1. CFD-modellin of Ni–FSF (2016) Unpublished Boliden Harjavalta internal report. In Finnish 2. Järvi J, Latostenmaa P, Osara K, Peuraniemi E (2014) Boliden Harjavalta smelter. In: Proceedings of 14th international flash smelting congress. Outotec, Espoo 3. Mäkinen T, Taskinen P (2006) The state-of-the art in nickel smelting: direct Outokumpu nickel technology. In: Sohn international symposium, vol 8. TMS, Warrendale (PA), pp 313–325 4. Suikkanen P, Haavanlammi K, Hietala K (2014) 20 years of DON process and its new benefits for Cu–Ni–PGM bulk concentrates. In: Proceedings of COM 2014, CIM 5. Taskinen P, Avarmaa K, Johto H, Latostenmaa P (2018) Fundamental process equilibria of base and trace elements in the DON smelting of various nickel concentrates. In: Proceedings of extraction 2018, CIM, SME & TMS

Redesign and Rebuild of the Pan Pacific Copper Flash Smelting Furnace Glenn Stevens, Tatsuya Motomura, Tomoya Kawasaki, Misha Mazhar and Gary Walters

Abstract After 40 years of operation with the original design, Pan Pacific Copper determined it was necessary to rebuild the Saganoseki Flash Smelting Furnace to continue safe operation. The original design employed a rigid steel frame, which, through hearth growth, led to severe distortion of the frame. Contributing to the continued growth of the hearth were thermal cycles that occurred during the government mandated annual shutdowns. Hatch designed a unique sprung bound, pivoting binding frame to maximize crucible size within the existing furnace footprint, while integrating the PPC designed cooling jackets. The bound system maintains tight brick joints, while the new conductive hearth design with integrated bottom cooling produces a protective freeze layer to accommodate higher furnace throughput. Minimization of furnace downtime for the rebuild was achieved through effective construction planning, highly trained contractors, and through an efficient start-up and ramp-up to full production. Keywords PPC



Hatch



Copper



Furnace



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Cooling

Introduction Pan Pacific Copper operates two copper smelters and three refineries in Japan; the Saganoseki Smelter and Refinery, the Hitachi Refinery, and the Tamano Smelter and Refinery of Hibi Kyodo Smelting Co., Ltd. The combined production capacity of the three smelters is 740,000 tonnes of copper per year, making it the largest copper producer in Japan.

G. Stevens (✉) ⋅ M. Mazhar ⋅ G. Walters Hatch Ltd., 2800 Speakman Drive, L5K 2R7 Mississauga, ON, Canada e-mail: [email protected] T. Motomura ⋅ T. Kawasaki Saganoseli Smelter & Refinery Pan Pacific Copper Co Ltd., 3-3382 Saganoseki, 879-2201 Oita, Japan © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_7

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Pan Pacific Copper’s Saganoseki Flash Furnace was originally built in 1973. The flash smelting furnace, located in Southern Japan, operates at 215 tonnes of feed per hour. After more than 40 years of operation with the original steel frame and much of the original hearth, Pan Pacific Copper (PPC) determined that it was necessary to modernize the flash furnace to continue safe operation. The original furnace settler design utilized a fixed steel frame which was unable to maintain tight brick joints during cool-down. This led to the formation of large gaps in the brickwork that would fill with molten products, especially upon reheat after a shutdown. The thermal cycling that leads to the formation and filling of gaps is referred to as ratcheting and is the primary cause of long term furnace growth. The ratcheting of the PPC furnace was exacerbated by government mandated annual shutdowns to inspect the waste heat boiler and the oxygen plant. The continued furnace ratcheting led to the lower endwall deforming by over 1 m and, as shown in Fig. 1, also led to furnace shell rupturing, and the distortion of the hearth support beams. Although 40 years of continued operation on the original hearth is considered exceptional by industrial standards, PPC decided to modernize the furnace to improve safety, increase efficiency, and reduce maintenance. Due to the recent trend of copper concentrate becoming higher in Sulfur and lower in Copper, PPC also acknowledged the necessity to upgrade the furnace cooling system. PPC partnered with Hatch for their ability to custom design a new furnace incorporating three dimensional bindings and forced bottom air cooling system while integrating PPC’s own patented copper cooler design. The sprung bound Hatch furnace design maintains tight brick joints throughout all stages of furnace operation while the new conductive hearth design promotes a protective layer on top of the hearth, all of which works to safely accommodate higher furnace throughput.

Fig. 1 Damaged furnace shell (left), damaged hearth beams (right)

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In addition to improving the overall furnace design, another important project objective was to minimize construction time. The Saganoseki Flash Furnace is the only primary smelting furnace at the facility and it was imperative to minimize furnace downtime. The construction management team of JX Engineering Co., Ltd. and Chiyoda Corporation was brought onto the project at the very early stages to ensure constructability was considered throughout the entire engineering and fabrication stages.

Furnace Design The 2017 rebuild of the PPC furnace was the first major rebuild in the furnace’s more than 40-year history. There were numerous upgrades incorporated into the new furnace design which improved safety, reduced required maintenance and worked to improve overall throughput: • Sprung bound frame to maintain tight brick joints throughout the life of the furnace and accommodate expansion and contraction cycling • Improved sidewall cooling jackets design to extend campaign life of the walls and accommodate a wider process variation • Increased slag retention time by increasing the overall furnace size • Reduced hearth thickness to reduce cost and construction time • Conductive hearth with bottom cooling to promote a frozen layer to accommodate higher heat load • Upgraded seismic design • Increased instrumentation for improved feedback and preventative maintenance planning A side by side comparison of the new and original furnace is shown in Fig. 2.

Binding System A successful furnace binding system applies compressive force in all three dimensions to maintain tight brick joints throughout all phases of the furnace operation. The binding system is particularly important during a cooldown phase as the refractory contracts, opening small gaps between bricks. As described earlier, without a properly designed binding system to close the gaps, thermal ratcheting of the brickwork can lead to excessive expansion.

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Fig. 2 Comparison of new furnace (Left) and original furnace (Right)

Horizontal Binding One of the challenges in applying the technical solution, and is common in most brownfield executions, is the space constraints. To maximize the furnace throughput, it was necessary to maximize the furnace size within the existing building footprint. In order to design a sprung bound furnace, while increasing the furnace volume, a unique pivoting binding design was created. A fixed cap beam was employed at the top of the binding frame which allowed for minimal clearances to fixed building steel. The binding springs at the base of the buckstay allowed the buckstay to rotate as the furnace expanded and contracted.

Vertical Binding To maintain tight horizontal brick joints, and maintain skew stability a novel hold-down system was incorporated into the design [1]. A crank in the sidewall shell captures the upper skew brick, while spring assemblies at the base of the shell maintains vertical pressure on the skew while providing a flexible support. The hold down system, improves hearth stability, minimizes space requirements and was designed such that the springs do not require adjustment during the full furnace campaign life.

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Seismic Design The Saganoseki smelter is located within an active seismic zone. To protect the furnace in the event of an earthquake, a series of seismic lock-up devices were installed to restrict the movement of the furnace and improve the containment of the bath in the event of an earthquake. During normal operation the devices allow the furnace to slowly expand and contract and lock-up during a seismic event. The lock-up devices work in conjunction with the fixed upper cap beam and utilize the refractory walls as shear walls to transfer the seismic load to the furnace foundations which are reinforced for the additional load. The design objective was to not apply any additional load to the building steel which would have required costly and time-consuming reinforcement to account for the additional loading.

Furnace Geometry The upgraded furnace design increased the bath volume by 16.7% through increasing the overall furnace size as shown in Table 1. Increasing the bath volume, in turn increases the slag retention time within the furnace while operating at the same feed rates. The original PPC furnace wall was designed at a 10-degree slope, as shown in Fig. 2. The original wall design was subject to considerable wall wear, as described in Sect. Wall Design. The slope of the wall assisted with maintaining the stability of the remaining wall once the refractory had worn away. The improved cooling jacket design within the furnace walls enabled the slope of the walls to be decreased to 2°. Reducing the wall slope allowed the bath area to be increased.

Hearth Design Prior to the 2017 rebuild, the PPC flash furnace was not fitted with bottom cooling. The hearth consisted of seven layers of arched refractory designed to insulate the bottom plate from heat. For this rebuild, PPC also decided to implement Hatch’s forced air cooling for the hearth to accommodate larger heat loads due to the change in concentrate grade. The upgraded hearth design included a thinner, more conductive,

Table 1 Comparison of original to new furnace size Variable

Original

New

Increase

% Increase

Furnace length Furnace width Furnace area at slag hole

20.7 m 7.2 m 149 m2

21.7 m 8.0 m 174 m2

1.0 m 0.8 m 25 m2

4.8 11.1 16.7

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refractory hearth combined with bottom cooling to reduce the hearth’s thermal resistance thereby increasing its cooling capabilities. The furnace refractory material selection was based on compatibility with the process conditions, thermal conductivity, mechanical properties, and resistance to hydration. The infill along with the hearth was selected to provide increased thermal conductivity compared to the existing furnace design. The increased conductivity raises the freeze isotherm of the matte and speiss higher into the upper hearth and promotes the formation of a protective build-up layer on top of the hearth. Material along the base course and infill perimeter was also selected to provide hydration resistance lining against the shell and bottom plate. Combustible expansion papers were inserted between refractory bricks to provide thermal expansion allowance during heat-up. The refractory expansion paper design is closely integrated with the furnace binding system to control furnace growth while ensuring constant and sufficient application of binding loads throughout the heat-up. Air cooling channels were integrated into the modularized furnace bottom girders for streamlined construction and installation. The twelve two-pass cooling system draws air into channels from one main duct located under the grillage beam assembly. Each channel inlet is fitted with a slide gate to balance flow through the system. Long and short thermocouples located throughout the furnace hearth were installed to provide comprehensive monitoring. These thermocouples are used to measure the temperature difference across a section of the hearth refractory to evaluate heat losses. Sacrificial thermocouples were also used during the furnace cold start-up to monitor the heating of the furnace hearth.

Wall Design The original water-cooled copper design utilized a vertical jacket, protected by a 350 mm thick refractory wall as shown on the left of Fig. 3. The photo on the right of Fig. 3 shows the condition of the refractory wall and cooler after 12 months of operation. The refractory wall in the slag zone has been eroded completely and the copper cooler is exposed to the furnace bath. PPC developed, and patented, a water-cooled jacket with copper fingers that extended into the furnace to retain the refractory bricks, as show in Fig. 4 [2, 3]. Following a 24-month campaign the majority of the refractory wall was still intact, resulting in a considerable thicker wall than achieved with the previous design. During the 2017 rebuild this cooler design was installed around the full furnace perimeter. In addition, the copper fingers provided much greater wall stability in the slag zone and gas phase of the settler [4, 5]. Moreover, PPC reviewed the heat load for each part of the settler, and the copper finger shape design was customized for each section of the furnace. As the result

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Fig. 3 Original water-cooled copper jacket design (12 month campaign)

Fig. 4 Redesigned water-cooled copper jackets (24 month campaign)

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of this, the appropriate cooling was achieved for each part of the furnace to maintain the proper thickness coating and the molten bath volume was increased (patent pending). Due to the addition of furnace bindings which allows the furnace to expand and contract, PPC incorporated lap joints and seal plates to accommodate the furnace movement while maintaining a tight joint between adjacent coolers which reduced the likelihood of molten leakage. The wall and roof interface was designed to provide allowance for relative movements, while maintaining a refractory layer able to insulate the steel shell and seal against the escape of freeboard gases. This was achieved by installing a combination of refractory ceramic fibre board and blanket, and managing brick joint locations.

Construction In order to minimize the downtime of the primary smelting furnace at Saganoseki, an aggressive 76-day schedule was prepared from feed off on September 21st, 2017. The furnace was rebuilt from the foundation up, including the complete furnace crucible and roof, reaction shaft transition coolers and roof, and the uptake shaft refractory. Overall, the schedule included 19 days for demolition and the removal of the frozen heal, 29 days for furnace steel construction, and refractory installation for 22 days. The furnace heat-up was on schedule, beginning on November 30th. Construction execution was as planned, with no delays to the overall project schedule. Feed on, and the first slag and matte taps were completed on schedule on December 6th, 2017. The success of the condensed construction schedule was influenced by project team cohesiveness developed and refined through feasibility, basic engineering, detailed engineering and construction planning phases. In addition, project team workshops, trial assemblies, modularization, sequencing of the construction, technical assistance, and highly trained and experienced contractors contributed to the construction success. The construction team consisted of design engineers, construction managers, construction supervisors, engineering managers, and quality control personnel from Pan Pacific Copper, JX Engineering, Chiyoda, and Hatch.

Workshops PPC, Hatch, and the construction management team (JX Engineering and Chiyoda) conducted multiple one to two-week workshops at the fabrication shop, discussing the details specified on the assembly drawings and the dependencies of each

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detailed activity within the steel and refractory schedule. Timing and sequencing of temporary support installation and removal, and installation scope boundaries were also discussed. It was critical to the success of the project and to the long campaign life expected from the new furnace that construction quality and tolerance were met without requiring rework or otherwise adversely impacting the construction schedule. Installation specifications created by Hatch provided requirements for achieving assembly tolerances, and controlling sources of error, and controlling installation quality.

Trial Assemblies An effective strategy of developing a confident construction schedule was the furnace steel trial assembly (Fig. 5) at the fabricator shop prior to shipment. Erection methodology, and means and devices to confirm and approve positional requirements were finalized during the trial assembly phase. Any minor interferences were also identified and corrected as required at the fabrication shop, saving the error from reaching site and impacting the shutdown schedule. Quality control requirements and allotted estimate times were assessed based on the time taken to execute during the trial assembly. Quality control supervisors and engineers that were scheduled to be on site from PPC, JX Engineering, Chiyoda, and Hatch were present at the time of trial assemblies. The effectiveness of the trial assembly added to the construction efficiency by applying lessons learned here to the construction execution. Select fabrication shop installation team members were also present on site for installation, which allowed efficiencies in construction execution.

Fig. 5 Furnace binding steel trial assembly

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Modularization During the engineering phase, a modularization plan was developed for furnace steel installation on site which was initially implemented during the trial assembly. The modularization plan included connecting, temporarily or as per design drawings where possible, the main binding frame components (buckstays, capbeam segments and waler beams). By connecting the binding frame components into four modules per sidewall, one per endwall, and four corner binding modules (Fig. 6) the number of lifts and movements were greatly reduced. The bottom girders were installed with the longitudinal and transverse tie rod segments installed through the bottom girder tie rod holes and seismic lock up devices were bolted to the bottom girders. The bottom girder module plan was implemented to ease installation due to limited access from underneath the furnace. Pre-assembly of grillage beam bolted connections, and endwall seismic lock-up devices attached to the grillage beams was also completed because of access constraints on site. Main binding spring set assemblies, and secondary spring sets were pre-assembled in the shop removing the need to mark, identify and install smaller components on site. The modularization improved the quality of the overall assembly, as modules were inspected in the fabrication shop and steel components were fixed to one another, minimizing the probability of installing a component incorrectly. Overall, the modularizations and preassembly planning effort was instrumental in achieving a shutdown period of 76 days.

Fig. 6 Binding frame components module (Left−Binding corner module, Right−Bottom girder and tie rod module)

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Installation System and Methods A skidding system was implemented on site to move furnace components into the furnace footprint, as shown in Fig. 7. The skidding system significantly reduced the number of heavy lifts required, improved safety and overall installation time. The load is kept low, which eliminates safety concerns with moving suspended loads. The skidding system required the removal of the existing tapping platform, and installation of a temporary platform for the tracks. Other attributes of the system included longitudinal and transverse tracks and temporary support beams. Once the modules or furnace steel components had been transported within the furnace footprint chain blocks on a rail system above were attached to pre-installed lifting lugs for final positioning.

Fig. 7 Skidding system (Left−Concept, Right−Grillage beam rails)

Fig. 8 Furnace heat-up curve

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Start-Up Following the successful completion of construction, gas burners were ignited on schedule on November 30th. As shown in Fig. 8, the actual heat-up of the refractory hearth followed the planned target very well. On December 6th the PPC furnace went from feed on at 9:25 am, to first slag tap at 6:25 pm, and first matte tap less than an hour later at 7:20 pm. Figure 9 shows the bath conditions 4.5 h after feed on. Within a week of first matte tap the furnace reached name plate capacity of 215t/hr of feed. After some consistent contraction of the furnace as a protective frozen heal was formed on the hearth, the furnace stabilized and reached an equilibrium state.

Closing Remarks The redesign and rebuild of the Saganoseki flash furnace has proven to be very successful to date. In the three months since start-up the furnace has been operating at an average feed rate above 200 tonnes of feed an hour, and on more than one occasion has exceeded the nameplate capacity of 215 t/hr. By utilizing the potential of the redesigned flash furnace, PPC is currently planning to increase the throughput further, to accommodate lowering copper grade in the concentrate.

Fig. 9 Bath conditions 4.5 h after feed on

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References 1. Hutchinson et al (2006) United States Patent No. 7,134,397 2. Yasuda Y, Motomura T (2009) Japanese Patent No. 2011-75183. Japanese Patent Tokyo 3. Yasuda Y, Motomura T (2009) Japanese Patent No. 2011075184. Japanese Patent Tokyo 4. Motomura T, Okamura H (2010). Japanese Patent No. 2012-57905. Japanese Patent Tokyo 5. Motomura T, Okamura H (2014) Japanese Patent No. 2015-197227. Japanese Patent Tokyo

Office, Office, Office, Office,

Modelling Metallurgical Furnaces— Making the Most of Modern Research and Development Techniques Evgueni Jak

Abstract Recent advances in analytical, experimental techniques, and computer-based theoretical modelling of fundamental properties and elemental processes, provide new opportunities to develop the next level of whole-of-reactor pyrometallurgical furnace models. These models have the potential to significantly improve the prediction of, and adding value to, industrial operations. In non-ferrous smelting, the starting point of these models is the development of multicomponent thermodynamic databases for gas-slag-matte-speiss-metal-solids phases supported by systematic experimental research. The whole-of-reactor-models additionally should take into account kinetic processes taking place at micro- and macro- scales, and other key factors. Examples of applications of the latest research tools and modelling approaches to analysis of industrial flash and top submerged lance (TSL) sulphide smelting processes are presented. Different levels of industrial modelling are discussed from elemental local reactions, through general and more detailed whole-of-reactor-models, to plant sections and further to whole plant operation models. Some principles for development of pyrometallurgical reactor models are discussed. Keywords Pyrometallurgy Thermodynamic modelling



Non-ferrous



Experimental

Introduction Statement Improving throughput, recovery and treatment of complex feed sources, decreasing inventories in pyrometallurgical operations can give significant potential benefits. These can be achieved by improved process stability, process and feed optimisation, the introduction of feed-forward process control and advanced process E. Jak (✉) PYROSEARCH, Pyrometallurgy Innovation Centre, The University of Queensland, Brisbane, QLD 4072, Australia e-mail: [email protected] © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_8

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planning. These improvements require accurate reliable modelling of pyrometallurgical reactors for information-based decision making. The present paper summarises some practical experiences in process modelling in pyrometallurgy with intention to highlight some of the issues commonly faced during implementation of fundamental thermodynamic modelling tools in industry.

Accurate Computer-Driven Models of Pyrometallurgical Reactors Are Needed to Achieve Significant Potential Improvements The operations of pyrometallurgical smelting, refining and recycling processes are complex as they involve control and adjustment of a large number of closely inter-related parameters. Many factors have to be taken into account including, but not limited to, the chemical partitioning and internal recycling of major and minor elements between phases (gas, slag, matte, metal) and multi-phase process streams, heat balance and temperature control, slag liquidus and proportion of solids, formation of frozen protective layers on the furnace walls and furnace integrity, off gas volume and temperature, vaporised elements and condensed dust carry over, waste heat boiler load. The relationships between the input and output parameters are complex, so too is the task of control and optimisation of the processes. The majority of pyrometallurgical processes, therefore, are currently operated using feedback control approach in which the input is adjusted based on the output only after the material has been processed in the reactor. The feedback approach (i) does not require absolute values leading to the lack of accurate data, and (ii) limits the possibility to improve control and to optimise operation. To improve process performance, to achieve feed and overall process optimization, and to implement feed-forward control it is necessary to be able to accurately predict the process outcomes as function of input control parameters.

Recent Advances in Analytical, Experimental and Computer Phase Equilibria/Thermodynamic Modelling Capabilities— Important Foundation for the Advances in Pyrometallurgical Process Modelling Recent significant advances in analytical, experimental and computer thermodynamic modelling provide important foundation for considerable progress in pyrometallurgy process modelling. Scanning Electron Microscopy (SEM) and Electron probe X-ray microanalysis (EPMA) enable advanced microstructural characterisation and accurate composition measurements at spatial resolutions of approximately 0.5 to ∼1–3 μ [1]. The

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recent development of the laser ablation inductively coupled plasma mass spectrometry (LA-ICPMS) microanalysis reduces the minimum elemental detection limit down to ppb at spatial resolution of approximately 20–80 μ [1]. Direct measurement of the compositions of phases in quenched samples eliminates the uncertainties related to the composition changes during equilibration and thus has enabled phase equilibria to be established over wide variety of systems not possible previously. Microanalysis of industrial quenched samples provides an exceptional opportunity to characterise processes taking place in the reactors. Increase in computer power, development of advanced theoretical models for complex solution phases and progress in thermodynamic database optimisations have resulted in the development of integrated commercial computer packages with extended, self-consistent, multi-component multi-phase databases, such as MTDATA [2], ThermoCalc [3], FactSage [4, 5]. These packages enable simultaneous predictions of all thermodynamic and phase equilibrium properties, such as the proportions and compositions of phases at given temperature and atmosphere, liquidus, solidus, thermodynamic activities of species, chemical partitioning of major and minor elements between condensed phases, vapour pressures, enthalpies of reactions and other key parameters essential in pyrometallurgy. In summary, the advances in (i) microanalysis analytical techniques, (ii) thermodynamic predictions and (iii) computer capabilities, now enable the development of powerful models of pyrometallurgical reactors that can facilitate significant, step-like improvements in the sector.

Introducing the Virtual Reactor and Pyro-GPS Computerised industrial reactor models for accurate prediction of the outcomes of the real industrial processes from input parameters and therefore for optimisation of processes are needed to facilitate significant improvements of pyrometallurgical operations; such models may be called Virtual Reactors. A Virtual Reactor model (Fig. 1) should be capable, for a given type of reactor, of predicting the temperature, amounts and detailed compositions of all phases in all output process streams as a function of all relevant input parameters over a wide range of conditions including input temperature, pressure, compositions, mineralogy of each input stream and other important thermochemical parameters. The models of the individual reactors can further be used within a broader computer-aided smart-decision-making systems of operation that may be called Pyro-GPS by analogy to GPS The Virtual Reactor and Pyro-GPS should incorporate a number of key components [6], including:

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Fig. 1 a Virtual reactor and Pyro-GPS concept, b components and stages of development and implementation (from [6])

1. Thermodynamic models to predict thermodynamic direction, driving force, extent and enthalpies of reactions (the latter defining the heat balance); and phase equilibria to describe the states (liquid, solid or gaseous), chemical compositions and proportions of the phases present in each of the process stream; 2. Physical property models to predict viscosities, surface tension, densities and other; 3. Micro-kinetics at a scale up to 20–100 μ taking into account the influence of the heterogeneous gas/solid/liquid reactions taking place in the pyrometallurgical reactor; 4. Macro-kinetic models to describe fluid flow and heat transfer at the full multi-meter reactor scale; 5. Plant data accuracy providing reliable information on thermochemistry of the all input and output streams including, but not limited to, temperature, composition, mineralogy; 6. Plant control accuracy to ensure stable operation at selected optimised conditions, and 7. Performance of the reactor within the overall multi-reactor flowsheet.

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Accurate prediction of thermodynamics and phase equilibria is an important foundation for any pyrometallurgical reactor model. The current status of development of the computerised thermodynamic packages with extensive multi-component multi-phase databases provide advanced predictive capability necessary for development of Virtual Reactor and Pyro-GPS. For example, Pyrometallurgy Innovation Centre (PYROSEARCH) at The University of Queensland supported by the consortia of leading international metallurgical companies is undertaking a research program on the high temperature copper-leadzinc-containing gas/slag/matte/metal/speiss/solids oxide and sulphide systems with major “Cu2O”–PbO–ZnO–FeO–Fe2O3–SiO2–S, slagging Al2O3, CaO, MgO and selected minor elements, including As, Sn, Sb, Bi, Ag, Au that utilises and developing further these recent capabilities. The research program involves integrated experimental and thermodynamic modelling studies using FactSage [7–9].

One-Step Reactor Model An example of the one-step reactor model of the Flash Smelting Furnace given in Fig. 2 (taken from Jak [10]) developed with the complex multi-component multi-phase thermodynamic database within computer package FactSage is given to demonstrate significant improvement in predictive capability of the pyrometallurgical process thermochemistry.

mechanical dust concentrate gas Coal or oil slag flux(SiO2) ma e O2+air heat loss

Fig. 2 One-step reactor model

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The reactor model takes as inputs the concentrations, temperatures and mineralogies of all feed streams including different concentrates, fuel (coal, oil or gas), recycled dust, flux and oxygen-enriched air. The materials remaining after the mechanical dust carry over are assumed to react producing gas, matte and slag with solids. The model was developed by (i) predicting equilibrium using complex thermodynamic models as a foundation (ii) with the kinetic calibration parameters (such as oxygen efficiency, heat balance, flux utilisation) introduced using well characterised synchronised input and output dataset. Predictions describe complex chemistry of all phases including ferric/ferrous ratio, copper and sulfur in slag, oxygen in matte, oxygen and sulphur partial pressures in gas, minor elements concentrations and partitioning between all streams, proportion of solids if any and composition of remaining liquid etc. Examples with detailed descriptions are given in a number of publications by the author [9–14]. These simplified thermodynamically-based models with the kinetic “calibration” parameters have important advantages for the modelling of the metallurgical reactors and is a significant step compared to the common solution of the heat and mass balance based on the use of spreadsheets with fixed partitioning coefficients. In particular, using this approach (i) the mass and energy balance is predicted together with (ii) the partitioning of the major and minor elements between all phases and (iii) phase equilibria (formation of solids, liquids and gas). Chemical reactions and the final phase compositions predicted using complex thermodynamic models of the multi-component solution phases 1. provide significantly more comprehensive descriptions of the complex chemistries of the output streams, 2. generate more accurate reliable interpolations and extrapolations 3. are valid over wider ranges of process conditions. These models can be very effective in different ways, for example as process advisers improving the control of the process, in the preparation of blends, in short and long term production planning, limited scale-up predictions, staff training. More sophisticated multi-step reactor models require more data and efforts to develop, and higher skills level to operate effectively.

Two-Step Counter-Current Reactor Model Example of a two-step counter-current reactor model application for more detailed analysis of material and energy recirculation in blast furnace is given to demonstrate that 1. complex processes can be predicted and 2. more advanced levels of process analysis can be achieved using models based on advanced thermodynamic databases. Predictions of alkali metals recirculation in iron blast furnace due to vaporization and re-condensation, and of associated energy transfer are important.

Modelling Metallurgical Furnaces—Making the Most of Modern … Fig. 3 Two-step iron blast furnace model

1Feed

10Off Gases

Gas Condenser

9- Condensed 8- Heat Exchange

109

2-Coke

Hearth Reactor

3- Heat Loss

4- Slag 7- Hot Gases 6-Air Blast

5-iron

The use of thermodynamic solution databases for slag is particularly significant when considering high alkali concentrations and extensive solutions that can occur in liquid and solid silicate phases. The iron blast furnace is a continuous, countercurrent reactor in which heat generated by the combustion of coke is transferred from the hot ascending gas phase to the descending condensed phases. The furnace is characterized by a thermal and chemical reserve zone in which gas and condensed phases approach equilibrium. In the following example (taken from Jak et al. [15, 16]) the iron blast furnace is described as a two-step countercurrent reactor model (see Fig. 3) consisting of Hearth Reactor and Gas Condenser. Coke and preheated air blast enter the Hearth Reactor. Fresh feed reacts with hot gases to form condensed material and heat exchange enthalpy (both predicted in the Gas Condenser) that enter the Hearth Reactor. Hearth Reactor predicts output streams including compositions, flow rates and temperature of the slag and liquid iron that leave the Blast Furnace; and of the hot gases that are passed into the Gas Condenser. Gas Condenser predicts the outcomes of a number of reactions including re-oxidation and condensation. Off gases leave the furnace, and condensed material and heat exchange enthalpy are passed into the Hearth Reactor. The calculations are organised in cycles and continued until the mass balance criteria are met. Predictions with complex thermodynamic models are used as foundation with kinetic factors introduced to describe real processes. The FactSage computer package macros and a thermodynamic database for the Al2O3–CaO–FeO–Fe2O3–Na2O–K2O–MgO–SiO2 system [15, 16] have been used for thermodynamic and phase equilibria predictions of the phases formed and the partitioning of all elements between the solid compounds, liquid slag, metal and gas phases in the Hearth Reactor and in the Gas Condenser.

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The model describes heat exchange and associated re-oxidation and condensation reactions in the gas phase during ascent of the gas as well as detailed compositions, phases and temperatures of all output and recirculated streams. Particularly useful is prediction and analysis of the effects of input parameters on recirculation of alkali metals and associated recycling of energy that takes into account thermodynamic properties of complex condensed solutions. More detailed description of the model and results is given elsewhere [15, 16]. This simplified model of the iron blast furnace demonstrates how important trends in furnace behaviour can be analysed. The same approach can be used to describe other counter-current reactors such as lead blast furnace, rotary kiln processes. This is a potentially powerful tool for analysis that can be further extended and refined.

Examples of Non-equilibrium and Kinetic Factors in Pyrometallurgical Furnaces Effect of Pyrometallurgical Reactor Type on the Partitioning of Major and Minor Elements Industrial data reported by Larouche [17] on partitioning of As, Pb, Sb and Bi between gas, slag and matte in copper smelting was analysed to identify trends [18]. Table 1 and Fig. 4 present the As partitioning for different types of reactors. The reactors are arranged (i) starting from suspension reactors (Outokumpu and Inco Table 1 As partitioning for different types of reactors [18] AVERAGE all furnaces MIN all furnaces MAX all furnaces ST. DEV. all furnaces AVERAGE Inco Flash Smelting Furnace (I.FSF) AVERAGE Outokumpu Flash Smelting Furnace (O.FSF) AVERAGE Mitsubishi S-furnace (M.S.) AVERAGE Reverberatory furnace (Rev.) AVERAGE Isasmelt (ISA) AVERAGE Teniente Converter (smelting furnace) (T.C.) AVERAGE Noranda process reactor (Nor. R.)

wt% slag/matte

% in matte

% in slag

% in gas

1.014 0.001 2.500 0.710 1.429

17.9 2.0 53.1 13.9 53.1

18.8 3.2 60.3 13.7 34.3

63.2 11.3 88.8 22.8 12.6

1.040

25.0

25.9

49.2

0.466

12.1

36.4

51.5

1.313

20.2

21.1

58.8

2.500 0.690

6.0 11.0

5.7 12.5

88.3 76.6

0.400

7.9

9.6

82.5

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gas

slag maƩe

Fig. 4 Partitioning of As between gas, slag and matte sorted by furnace type. In the labels, last two numbers mean final wt%Cu in matte and %O2 air enrichment, respectively. Black lines show industrial average [18]

flash furnaces) where concentrate reacts in the gas stream (ii) to the bath furnaces (TSL, Teniente and Noranda reactors) where O2-enriched air is blown into the molten bath and concentrate is added and reacts in the molten agitated bath. Figure 4 demonstrates that the As partitioning to the gas increases from suspension reactors to the bath type furnaces. This trend (if not affected by data accuracy and recirculation practices) is related to the importance of the micro- and macro-kinetic of the processes that depend on the furnace type when other factors are the same. Strong effect of the type of pyrometallurgical reactor on partitioning of minor elements demonstrated in this section highlights the importance of non-equilibrium factors in pyrometallurgical processes. Present example also demonstrates importance of the accurate plant data.

Example of Identification of Elemental Reactions in TSL Bath Smelting SEM and EPMA analyses of the slag sample quenched from the copper TSL furnace presented in Fig. 5 taken from Henao et al. [19] indicate that the slag contained small spinel crystals and suspended matte droplets. The composition of the matte droplets measured with EPMA was found to be close to the bulk matte composition indicating conditions close to the equilibrium between slag and matte in this well-agitated TSL process—this fact established by the microanalysis of the plant sample is important for the development of the TSL model.

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Fig. 5 Micrograph of the slag sample quenched from Cu TSL reactor (Taken from Henao et al. [19])

The matte droplets have a high-copper rim—this is consistent with the oxygen transfer mechanism from the oxygen-enriched air bubbles introduced through the lance to the sulphide feed through the Fe2+/Fe3+ couple in slag—from 0.5O2,gas + 2FeOslag = Fe2O3,slag to 3Fe2O3,slag + FeSmatte = 7FeOslag + SO2,gas. The ability to detect elemental reactions with microanalysis tools provides new opportunities to quantitatively investigate the reactions and assist in developing models describing processes in real industrial reactors.

Examples of Analysis of Elemental Reactions in the Flash Furnace Following example demonstrates the application of advanced tools to analyse the elemental processes in the industrial flash smelting furnace. Slag samples were quenched from several locations along the flash and electric furnace starting from 1. tapped matte 2. the matte-slag interface in the middle of settler, 3. top slag layer in the middle of settler, 4. slag tapped from the flash furnace, and 5. slag tapped from the electric cleaning furnace. The microstructures of the samples were analysed with SEM and compositions of phases were measured with EPMA. Figure 6 presents copper concentration in matte measured with EPMA. The matte grade in the suspended matte droplets increased from the slag-matte interface, through the middle of the settler to the tapped slag and further to the slag from electric cleaning furnace, and were all higher than the matte grade in the bulk matte. These results indicate that the process conditions vary with locations in the flash furnace; this information is essential for the development of a more accurate model of this process.

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Fig. 6 Variation of matte in different locations of a flash furnace

Multi-step Model of the Flash Furnace A one-step reactor approaching equilibrium is a possible first approximation of some pyrometallurgical processes. Any pyrometallurgical process, however, represents a combination of areas with different reactions and different conditions in various parts (or locations) of the reactor. For example, analysis of the slags from different locations of the flash smelter presented above indicates that different conditions are present in that reactor. Different reactions taking place in different locations in a copper Flash Smelting Furnace are illustrated in Fig. 7 [10]. The flash smelter may be viewed as a multi-stage reactor. I. Reaction shaft: The oxidation of the sulphide minerals takes place in the Reaction Shaft. Part of the feed (surface of larger particles and whole small particles) is over-oxidised so that the excess of oxygen in the condensed form is produced (e.g. ferric iron Fe2O3, CuO, Cu2O, CuSO4), at the same time, the core of larger particles may remain un-reacted due to kinetic limitations and stay under-oxidised (e.g. FeS, Cu2S etc.). The over- and under-oxidised particles do not practically react with each other in the reaction shaft due to the low chances of coalescence and relatively slow diffusion of main reactants through the gas phase. The formation of the solid spinel Fe3O4 or Cu2SO4 rim inhibits reactions with the gas.

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Fig. 7 Schematic diagram illustrating possible reactions in the Outokumpu copper flash smelting Furnace

II. Settler beneath the Reaction Shaft: After most of the material from the reaction shaft (including over- and under-oxidised particles) is plunged into the bath underneath the Reaction Shaft, all particles are forced in direct contact with molten slag and matte resulting in the establishment of condensed phase interactions, reactions, convective and diffusive mass transfer etc. The condensed oxygen from the over-oxidised particles will react with under-oxidised material including un-reacted sulphides thus producing SO2 gas. Description of some of these reactions and further references can be found in Taskinen [21]. III. Settler far from Reaction Shaft: After most of the oxidation/reduction reactions are completed in the settler area close to the Reaction Shaft, separation of the matte droplets from the slag layer takes place in the rest of the settler. Some further particles may still fall down onto the slag from the gas area above the settler, some further reactions may take place between the gas and slag phases and further down in the bath—the occurrence and extent of these are expected to be significantly less important than the reactions in and underneath the Reaction Shaft.

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Concluding Statement In conclusion, the above examples demonstrate that basic fundamental equilibrium models alone are not sufficient to accurately describe real industrial processes— many non-equilibrium and other factors must be incorporated. Characterising and taking non-equilibrium and other factors in real processes is essential and requires (i) special targeted applied fundamental research programs to characterise micro-kinetic and macro-kinetic processes in real furnaces, and (ii) special programs to improve the accuracy of the plant data and process control.

Possible Steps Towards the Development of the More Accurate Pyrometallurgy Reactor Models Snap—Shot Campaigns—Possible Step to Overcome Plant Data Accuracy Limitations Many pyrometallurgical furnaces are operated using feed-back control approach that is based on the gradual adjustment of the input control parameters relative to the previous values based on the observation of the output after the material passed through the reactor. The feed-back control therefore does not require accurate absolute data on the chemistry and temperature of the input and output streams. That results in the lack of the reliable data necessary to develop reactor models. To overcome this limitation, specially targeted snap-shots campaigns can be done in conjunction with independent laboratory measurements and modelling of the input—output results. For example, the following synchronised collection of operational plant data together with additional temperature measurements and sampling are can be undertaken for a flash furnace: 1. Routine operational plant data: collection of all “normal” input and output operational plant data. The continuous data from plant data management system such as PI (e.g. feed rates) averaged over a designated periods of time should be synchronised with the fixed-time data (e.g. routine sample collection or dip thermocouple temperature measurement). Synchronisation should take into account residence time on conveyors and in the furnace. All input and output streams should be incorporated, including concentrate, dust, coal or oil, air, oxygen, slag, matte etc. Temperatures, compositions and mineralogies of the streams are important. The number of locations included into the campaign can be selected depending on the targeted level of model. For example, the key input and output streams (feed A, tapped matte J, tapped slag F, and dust I) can provide an adequate dataset.

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Fig. 8 Possible flash furnace sampling points for a snap-shot campaign: A—feed sample; B— observation hole in the reaction shaft; C—access point on settler roof after reaction shaft; D— access point in the middle of the settler roof; E—access point for bath level measurements after uptake shaft; F—tapped slag; G—granulated slag; H—accretion on the uptake shaft; I—dust at the exit from uptake shaft; J—tapped matte

2. Additional Temperature Measurements by dip thermocouples in the selected points (for example see Fig. 8), including point E dip into slag (middle of the slag layer), point F—tapping slag, point J—tapping matte. 3. Additional control plant samples (synchronised with additional measurements of operational characteristics and with routine samples assays) may be taken, including feed and dust samples at point A, dip slag sample at point E, tap slag sample—point F, quenched matte sample—point J. The samples and data collection should be done only during normal stable operation of the furnace in the middle or end of the bed; unstable periods at the start of the bed or after stopping of the feed due to operational issues must be avoided. All samples bulk compositions should be analysed at the plant as well as an independent laboratory. The results of such studies have advantage of (i) eliminating plant data uncertainties, (ii) provide complete input-output set necessary for developing/calibrating reactor models. Detailed characterisation of the process can assist to take into account kinetic factors and enable to make further informed decisions.

Improved Analytical Capability for Characterisation of Minor Elements Distribution in Industrial Samples The next example [1] demonstrates recently developed capability to analyse low concentrations of minor elements; this is particularly important for hazardous and

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precious metals especially for the treatment of complex feed sources. Using the combination of EPMA and LA-ICPMS the minimum detection limit of elements can be reduced down to ppb (see Table 1). The development and application of the EPMA + LA-ICPMS techniques were used to analyse the laboratory and industrial quenched slag samples (see Fig. 9). The results for the industrial Cu smelting slag samples are given in Fig. 10. The advanced analytical techniques provide high accuracy data necessary for development of models and for informed evidence-based decision making.

Fig. 9 Microstructures of as-received smelting slag (left) and electric furnace slag (right)

Fig. 10 Distribution coefficients of minor elements between slag and matte measured with EPMA and LA-ICPMS in industrial an electric furnace slags [1]

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Detailed Analysis of the Minor Elements Partitioning Between Streams The combination of the snap-shot campaign and the one-step thermodynamic model predictions was used to analyse thermochemistry of the flash furnace including major and minor elements partitioning [20]. The recirculation dust in the furnace generated by (i) mechanical carry over and (ii) vaporisation/recondensation process was analysed. From the data on dust input, the recirculation of different elements was calculated as: Recirculation fraction RF = ðelement in dust tphÞ ̸ ðelement in ``fresh'' feed tphÞ. The overall average recirculation fraction for all elements was estimated. The recirculation values of individual elements relative to the average recirculation factor are given in Table 2. The recirculation fraction RF values for Cu, Fe, SiO2, CaO, MgO and Al2O3 were found to be close to the overall average recirculation fraction and were assumed to have just mechanical carry-over component. The recirculation fraction for S was below the average RF reflecting reaction of S to SO2 taken into the gas stream, and some sulphur returned with the dust was assumed to be in the form of sulphates. The recirculation fractions RF for volatile elements Pb, Zn, As, Sb and Bi were above the average RF indicating that the vaporization/ recondensation cycle for these elements is significant. The one-step model of the copper smelting process was developed as described in the above sections with reference to Fig. 8. The FactSage computer package was used together with the thermodynamic database for the multi-phase system with major “Cu2O”–PbO–ZnO–FeO–Fe2O3–SiO2–S, slagging Al2O3, CaO, MgO and selected minor elements, including As, Sn, Sb, Bi, Ag, Au developed at Pyrosearch

Table 2 Detection limits of EPMA, LA-ICPMS, XRF and ICP-OES [ppm] [1] Analysis techniques Elements

EPMA

LA-ICP-MS

BULK

As 185 0.11 100(XRF) Pb 288 0.19 100(XRF) Zn 199 0.88 100(XRF) Sn 318 0.09 50 (ICP-OES) Sb 281 0.02 10 (ICP-OES) Bi 192 0.02 10 (ICP-OES) Ag 122 0.01 1 (ICP-OES) Au 191 0.01 0.01 (ICP-MS) Note XRF: X-Ray Fluorescence analysis; ICP-OES: Inductively Coupled Plasma—Optical Emission Spectrometry; the reported detection limits of EPMA and LA-ICP-MS were based on the measurement results on the slag phase only; the detection limits for the bulk chemical methods were provided by Australian Laboratory Service

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Table 3 Comparison of plant and predicted values of the recirculated fraction of elements relative to the average recirculaiton Total

Cu

S

Fe

SiO2

CaO

MgO

Al2O3

Pb

Zn

As

Sb

Bi

Au

Ag

Recirculated fraction relative to the average

1.14

0.33

0.91

0.56

0.73

0.97

0.86

4.83

1.93

19.6

2.74

15.63

0.69

1.42

Model predictions

1.00

0.89

1.00

1.01

1.00

1.00

1.00

3.33

1.68

2.06

16.63

1.00

1.04

8.44

[7–9]. The partitioning of major and minor elements between gas, slag and matte process streams was predicted. Table 3 gives comparisons of the partitioning of minor elements from the plant data and the values predicted using this one-step reactor model. The predicted recirculation fractions for Cu, Fe, SiO2, CaO, MgO and Al2O3 were found to be close to the plant data and are all around 1 indicating a low degree of vaporization/ recondensation component. The recirculation value for Pb, Zn, As, Sb and Bi are all above 1 indicating significant vaporisation. The predicted recirculation values for Pb, Zn, Sb and Bi are close to the plant data, the value for As is less than half the plant value—this may be an indication of (i) importance of the kinetics for As, (ii) uncertainties of the plant data due to the addition of other unaccounted As-rich streams into the dust as well as (iii) uncertainties of the thermodynamic model. Important conclusions from this example are that snap-shot sampling campaigns complemented by predictions with one-step model using advanced thermodynamic database 1. remove a many uncertainties, 2. provide foundation for significantly more reliable process analysis and optimisation, 3. help to identify discrepancies for further analysis and improvements, and therefore 4. enable further model improvements to be focused on other issues to achieve even better quantitative predictions. The predictions demonstrate that non-equilibrium kinetic and other factors are important and must be taken into account during model development.

Summary Statement on the Kinetics and Other Factors in Pyrometallurgical Furnaces The improved level of knowledge as well as the advanced level of the research tools (analytical, experimental and theoretical computer modelling tools) now provide new opportunities in detailed and quantitative investigation of the complex reactions taking place in the industrial furnaces; this can greatly assist in the identification of the limits in optimisation of the processes, help to understand the reasons for some operation instability and therefore increase availability and decrease costs of the operations.

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Different Level Models Different levels of model sophistication (i) require different level of efforts/costs to be developed, and (ii) can provide different levels of information and accuracy. The selection of appropriate level of model is important. The following model levels may be identified: i. local reactions

ii. one step reactor model

iii. multi-step reactor models

iv. whole reactor CFD,

v. multi-unit plant sections, large plant sections or whole plant

• focusing on a particular part of the reactor or a particular property; e.g. “fluxing” model to provide guide to the operators/metallurgists—can be given as a rule e.g. “free silica” equation, table, fluxing diagram • describing whole reactor including mass and energy balance, and phase equilibria described with thermodynamic model and kinetic factors approximately described with a limited number of kinetic “calibration” parameters—these types of models can be used for process control as advisers, for design and feasibility studies • describing whole reactor as a combination of several one-step “sub-reactors” (sections, components) such as 2-step reactor-condenser model for counter-current reactors [BF reference]; each sub-reactor can be modelled with thermodynamic model and kinetic factors described by the special model parameters and the sub-reactors linked by the connecting mass and enthalpy streams • describing whole reactor using CFD modelling packages (e.g. Fluent [22] etc.); connection of sophisticated thermodynamic/phase equilibria models is usually limited due to the computation time, data-transfer/handling complexity and convergence limitations of the multi-component multi-phase thermodynamic models with complex solutions; the “architecture” for these models may involve (a) development of simplified phenomenological model, e.g. a polynomial description with limited accuracy and importantly—limited range of applicability using sophisticated thermodynamic/kinetic models as a basis, followed up by b) incorporation of the simplified phenomenological models into the CFD model (a) simple sections without recycling streams may be described using the multi-component multi-phase complex solution thermodynamic models as a basis for each operational unit—similar to the multi-unit reactor (b) the plant sections with recycling streams and large plant sections or whole plant may be modelled, similar to CFD, using two steps—1. developing simplified polynomial description of each operational unit using complex thermodynamically-based models, and 2. using adequate flowsheet packages (e.g. ASPEN [23], SysCad [24], MetSim [25]) for the whole plant section with the simple phenomenological polynomial descriptions of each operational unit

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It is essential to identify and make clear the correlation between (i) the model use, (ii) required complexity, accuracy and the range of applicability, and (iii) required data, sampling, development and implementation efforts needed to achieve targeted functionality and use of the model.

Some Modelling Principles The Model Development Versus Model Validation As discussed above, the prediction of the output from the known input for any pyrometallurgical reactor requires incorporation of a number of key components including thermodynamic and phase equilibria model, micro-kinetic, macro-kinetic, industrial plant data accuracy and industrial plant control accuracy. However, it is common that the direct use of the thermodynamic model without any assessment of the micro-kinetic, macro-kinetic and industrial plant data accuracy is referred to as as “model validation” –this is inadequate approach. The description of the link between input and output is actually the real process model development and calibration that should adequately describe all steps between thermodynamics and the process.

Furnace Modelling Approach as a System of Equations— Importance of Appropriate System of Input and Output Parameters for Model Development and Further Use Many interrelated factors simultaneously affect different outputs of the pyrometallurgical reactors as a function of several input control parameters. One input parameter usually affects several output parameters. Each output parameter can be influenced by several input parameters. It is not usually possible to identify direct link of one to another. In this multi-factor environment it is essential to identify the input/output couples with the most sensitive link. In the control as well as in modelling and process optimisation it is essential to identify the adequate system with the same number of input and output parameters —similar to the system of equations where the number of unknowns should be equal to the number of equations to have a single solution. In case of process optimisation—for each target output it is essential to identify a particular input parameter. In case of process model development—each model calibration parameter must have sufficient data to be adequately determined by the reliable industrial data. Too complex and sophisticated models with too many adjustable parameters can result in significant uncertainties in the process description since (a) many “solutions” are possible and (b) the values of the model calibration parameters cannot be defined. It is essential then 1. to select the level of model

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complexity that corresponds to the available plant data to be adequately calibrated, and 2. to ensure sufficient data by accuracy and by type are collected to identify model parameters through the synchronised data collection campaigns such as “snap-shot” described above, fundamental research of processes in the reactors and other. This represents one of the difficulties in modelling, control and optimisation of the pyrometallurgical reactors.

Example of a Simplified Three-Input-Output System of Model Parameters In any pyrometallurgical process the three key output target parameters are always essential to control operation (see Fig. 11) • Chemical targets—partitioning of elements between phases, • Heat balance, and • Phase equilibria, among many others, such as off gas volume and composition, minor elements partitioning and recycling, waste heat boiler load. Since all output parameters are influenced simultaneously by each input parameter and the effects are complex, the appropriate kinetic “calibration” parameter can then be selected for each output target at the model development stage. For example, as illustrated in Fig. 11, the oxygen efficiency, heat loss and flux utilisation parameters can be introduced and derived from the well established plant data set with known output matte grade, temperature and slag composition. The model then is developed by 1. taking complex thermodynamic model as a foundation, and then development of the kinetic “calibration” parameters for each couple of the key input—output parameters describing kinetic and other non-equilibrium factors using well characterised synchronised input and output dataset, these included oxygen efficiency, heat losses and flux utilisation. The

Main control input pyrometallurgical reactor Kine c parameters “calibra on” parameter Oxygen coefficient (O 2/feed ra o) Oxygen Dura on of blow and rate efficiency O2 enrichment, fuel (coal, oil), dust and reverts, feed rate, composi on (e.g. Heal loss Cu/S/Fe) and mineralogy, electric power, SiO2 and CaO fluxes, slagging impuri es Flux levels (e.g. Al2O3, CaO, MgO) u lisa on

Target output parameters

condi ons

/

Chemical - e.g. ma e grade, Cu or Pb in slag, %S in blister Heat Balance - temperature

Phase equilibria - Liquidus, % solids, freeze-lining thickness, frequently controlled by Fe/SiO2, CaO/SiO2 in slag

Fig. 11 Input and output parameters in pyrometallurgical reactors

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oxygen efficiency, heat balance and flux utilisation can be established from the well-characterised accurate dataset on output matte grade, temperature and slag composition respectively. The overall oxygen efficiency takes into account the extent of reactions in the furnace and air ingress due to the draft. The heat loss is calibrated initially using available reliable input/output data and temperatures to take into account not only the actual heat losses due to radiation, water cooling etc., but also the systematic model uncertainties. These model parameters can then be used for the predictions. Introduction of more input–output parameters would require to obtain more ploant data respectively. More sophisticated models would require more data to develop model parameters. The important point from this section to be highlighted is importance of adequate level of model sophistication directly dependent on the data available for the model development: no data—no parameters; if more parameters are needed—the data should be obtained. There should not be a model validation approach, there should be model development and calibration.

Conclusions The recent advances in analytical, experimental and modelling capabilities for the pyrometallurgical processes have provided new opportunities to model and therefore to optimise and improve the operations. Sophisticated models can result in significant improvement of the process financial and operational outcomes that justify the investment. The implementation of these advanced tools requires systematic approaches that take into account the correlation between model sophistication and efforts required for development of the model, as well as the correlation between model sophistication and the profits that can be generated through implementation. Acknowledgements The author would like to thank many industrial sponsors and many R&D and operation staff in the companies for the financial and technical support The author would like to acknowledge help of colleagues in preparation of this paper including Dr. Shishin, Dr. Hidayat, Prof. Hayes and others. Special acknowledgement and thanks to Prof. Hayes for the support and significant input over many decades into the research and education in the metallurgy sector.

References 1. Chen J, Allen CM, Azekenov T, Ushkov L, Hayes PC, Jak E (2016) Quantitative determination of partitioning of trace/ultra trace elements between slag and matte generated in copper smelting process using microanalysis techniques, Copper 2016, Kobe, Japan, November 2016 2. MTDATA: Teddington, UK. www.npl.co.uk

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3. Thermo-Calc: Stockholm, Sweden. www.thermocalc.com 4. FactSage: Montreal, Canada. www.factsage.com 5. Bale CW, Belisle E, Chartrand P, Decterov SA, Eriksson G, Gheribi AE, Hack K, IH Jung YB, Melancon KJ, Pelton AD, Petersen S, Robelin C, Sangster J, Spencer P, Van Ende MA (2016) FactSage thermochemical software and databases, 2010–2016, CALPHAD 54: 35–53 6. Jak E (2018) The role of research in pyrometallurgy technology development—from fundamentals to process improvements—part 2—future opportunities. In: Peter Hayes symposium on pyrometallurgy, Extraction 2018, Ottawa 7. Hidayat T, Hayes PCC, Jak E (2018) Microanalysis and experimental techniques for the determination of multicomponent phase equilibria for non-ferrous smelting and recycling systems. In: Peter Hayes symposium on pyrometallurgical processing, Extraction 2018, Ottawa, Canada 8. Shevchenko M, Hayes PC, Jak E (2018) Development of a thermodynamic database for the multicomponent PbO–“Cu2O”–FeO–Fe2O3–ZnO–CaO–SiO2 system for pyrometallurgical smelting and recycling. In: Peter Hayes symposium on pyrometallurgical processing, Extraction 2018, Ottawa, Canada 9. Shishin D, Hayes PC, Jak E (2018) Multicomponent thermodynamic databases for complex non-ferrous pyrometallurgical processes. In: Peter Hayes symposium on pyrometallurgical processing, Extraction 2018, Ottawa, Canada 10. Jak E (2012) Integrated experimental and thermodynamic modelling research methodology for copper and other metallurgical slags. In: The 9th international conference on molten slags, fluxes and salts, Keynote Invited Lecture, Molten 12, Beijing, China, May 2012, paper w77 11. Nikolic S, Shishin D, Hayes PC, Jak E (2018) Case study on the application of research to operations—calcium ferrite slags. In: 7th International symposium on advances in sulfide smelting, Extraction 2018, Ottawa, Canada 12. Shishin D, Hidayat T, Decterov S, Belov GV, Jak E (2016) Thermodynamic database for pyrometallurgical copper extraction, Copper 2016, Kobe, Japan, November 2016 13. Hawker W, Vaughan J, Jak E, Hayes PC (2016) The synergistic copper process—a new process route for low energy copper production, Copper 2016, Kobe, Japan, November 2016 14. Shishin D, Hidayat T, Decterov S, Jak E (2016) Thermodynamic modelling of liquid slag-matte-metal equilibria applied to the simulation of the Pierce-Smith converter. In: 10th International conference on molten slags, fluxes and salts, Molten 2016, May 2016, Seattle, Washington, USA, pp 1379–1388 15. Jak E, Hayes P, Pelton AD, Decterov SA (2009) Thermodynamic modelling of the Al2O3– CaO–FeO–Fe2O3–PbO–SiO2–ZnO system with addition of K and Na with metallurgical applications. In: 8th International conference on molten slags, fluxes and salts, January 2009, Santiago, Chile, pp 473–490 16. Jak E, Hayes PC (2012) The use of thermodynamic modeling to examine alkali recirculation in the iron blast furnace. High Temp Proc Mat Proc 31(4–5):657–665 17. Larouche P (2001) Minor elements in copper smelting and electrorefining. McGill University, Montreal. Available from 22077 18. Shishin D, Jak E (2015) Report on the minor elements distribution in copper smelting systems. In: Private communications, pyrosearch, The University of Queensland 19. Henao HM, Ushkov LA, Jak E (2012) Thermodynamic predictions and experimental investigation of slag liquidus and minor element partitioning between slag and matte in support of the copper Isasmelt smelting process commissioning and optimisation at Kazzinc. In: The 9th international conference on molten slags, fluxes and salts, Molten 12, Beijing, China, May 2012, paper w78 20. Shishin D (2018) Private communications, pyrosearch, The University of Queensland

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21. Taskinen P (2012) Direct-to-blister smelting of copper concentrates: the fluxing chemistry. In: Ragnarsson L, Jonsson P (ed) The Seetharaman Seminar 2010, Department of Materials Science and Engineering, Royal Institute of Technology, KTH, Sweden, pp 110–120 22. ANSYS Fluent: Caninsburg, PA, USA. www.ansys.com 23. Aspen Plus, Bedfort, MA, USA. www.aspentech.com 24. SysCAD, Perth, Australia. www.syscad.net 25. Proware, Tucson, AZ, USA. www.metsim.com

Pyrometallurgical Processing of Desulphurization Slags Christoph Pichler, Jürgen Antrekowitsch and Karl Pilz

Abstract A special desulphurization slag accrues during the processing of iron ore to steel, which is needed to generate high quality steel. Sulphur removing is done by different technologies, but the most common one is the generation of a high sulphur containing slag by sulphur affine elements. Currently, this desulphurization slag is partly recycled, however, huge amounts are still dumped. To utilize this desulphurization slag for future purposes, a recycling process must be established. Therefore, first characterization of this material started some years ago at the Chair of Nonferrous Metallurgy, Montanuniversität Leoben. Subsequent trials in lab scale size were performed, also some investigations for different treatment steps were evaluated, starting with a parameter study in a hot stage microscope at various gas-atmospheres. Finally, the developed recycling process concept was verified in pilot scale trials to confirm the new recycling technology. Due to the successful treatment process, also a patent was applied. Keywords Recycling



Desulphurization slag



Landfill

Introduction Due to the strong increasing crude steel production, the occurring by-products and residues from the iron- and steel industry are rising as well. This fact in conjunction with an environmental aspect is crucial to intensify the investigation in the field of recycling wastes and by-products from the iron- and steel industry. Some successful applications of utilizing materials from this industry sector are already in operation like the standardized cement from blast furnace slag because of its latent-hydraulic properties [1, 2]. Furthermore, the basic oxygen furnace slag can be used for road C. Pichler ⋅ J. Antrekowitsch (✉) Chair of Nonferrous Metallurgy, Montanuniversitaet Leoben, Leoben, Austria e-mail: [email protected] K. Pilz Voestalpine Stahl GmbH, Linz, Austria © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_9

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construction in different layers, due to its hardness [3]. Some other by-products cannot be used directly in a further application, because of their properties or chemical composition. One example is the desulphurization slag, which is the product at hot metal desulphurization. Such a sulphur removal is mandatory in the iron and steel industry, because sulphur is well known as steel parasite [4, 5].

Sulphur in the Iron- and Steel Industry This short chapter describes the influence of sulphur in the steel product and the possibilities for a successful removing.

Influence of Sulphur and Its Input Material Beside phosphor, sulphur is one of the major steel parasites. The complete dissolution of sulphur leads to the formation of iron sulphide. At fast cooling rates, the iron sulphide separates at the grain boundaries. This compound builds together with iron a eutectic with low melting temperature, which is responsible for the red

Fig. 1 Binary system of sulphur and iron [6]

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shortness during hot forming. The iron sulphide itself leads to the hot shortness at about 1200 °C. The binary system of sulphur an iron is pictured in Fig. 1 [4, 5]. Manganese forms together with sulphur a compound, which leads to problems during welding processes, and lamellar fractures occur. Only at some special steel types the sulphur is needed, like free-cutting steels. In earlier times lead was used in such steel types, which was prohibited in the last years and was substituted by sulphur, since of similar properties. Typically, sulphur contents are 0.15–0.3% in free-cutting steels [4, 5, 7]. The main source for sulphur in hot metal is the used coke at the blast furnace process. About 80% of the total sulphur input is contributed by this material. The remaining part of the sulphur feed is coming from ores and additives [5, 8].

Desulphurization Technologies for Hot Metal Different methods can be used for a successful desulphurization. Examples are the removal via the gas phase, through diffusion equalization, with sulphur affine metals or the usage of the partition equilibrium between a slag and iron. Within the hot metal or steel production different desulphurization technologies take place during different processing steps. The first removal is done in the blast furnace itself by the slag, where about 4% of the fed sulphur remains in the pig iron. The main step is done between the blast furnace and the steel plant in a pig iron ladle or torpedo ladle. To reach the defined level, the secondary metallurgy of steel is the last possibility to perform a desulphurization [5]. A summary of possible concepts for the desulphurization, divided into groups, is given in the following Table 1. Depending on the operated desulphurization process various amounts of desulphurization additives are needed and also the required time together with a temperature loss is varying [5].

Table 1 Overview of possible desulphurization concepts an additives Group

Process

Place of treatment

Type of desulphurization agent

casting stream mixing process Mechanical mixing process Pneumatic mixing process Electromagnetically mixer

soda desulphurization

Ladle/mixer

Soda

Diverse mixer; Hoesch mixing process Injection lance; injection process Electrochemical gutter

blast furnace gutter; pig iron ladle Torpedo ladle

CaO; CaC2; soda/ lime (1:1) CaC2 + CaO; Mg and Mg + CaO Pre-melted slag

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Fundamental of the Roasting for Sulphur Removing For special types of desulphurization slags some processes are available [9, 10], but in most cases a successful concept is not available and therefore it gets dumped. This information is also mentioned in the IRC Reference Report Best Available Techniques (BAT) Reference Document for Iron and Steel Production from March 2012. One main idea is the separation of sulphur without generating of a new waste. This can be achieved through a removal of sulphur in the off-gas linked to a sulphuric acid generation. In the field of metallurgy, the roasting process is typically used to generate metal oxides from sulphide ores. This leads to the fact that such a roasting step can be used to treat sulphur containing slags as well. Different roasting technologies are available like the multiple hearth furnace, the fluidized bed roaster, a rotary hearth furnace, a shaft furnace or a sinter belt. All these processes are working in solid state [11–14].

Recycling of Desulphurization Slag from the Pig Iron Desulphurization The considered desulphurization slag is generated in a hot metal ladle desulphurization through an injection lance. The obtained slag contains the sulphur and gets removed from the ladle. Here, a separation in three different grain size fraction can be achieved. The biggest fraction are so called “skulls” (>120 mm) which can be used as scrap in the basic oxygen furnace or an electric arc furnace. With respect to the sulphur content, in the most cases it is possible to recycle the “middle” sized-fraction (10–120 mm) in the blast furnace, but too high sulphur levels avoids such a reuse. In this case the middle fraction has to be treated together with the “fine” fraction (0–10 mm), which has the highest sulphur content. At the moment, no satisfying concepts are available to remove the sulphur from these fractions. Therefore, intensive research in the past leads to the following described process design.

Characterization and Parameter Definition in a Hot Stage Microscope (Lab Scale Trials) For this investigation the fine fraction of the desulphurization slag was used, that typically contains 4–6% S, CaO, MgO, C, SiO2, FeO, Fe2O3 and Femetallic. The first step was a detailed chemical analysis of the used material to obtain information about the detailed composition.

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To clarify the possibility for a roasting process, a hot stage microscope was applied. Usually such a device is used to determine the softening behaviour of slags, dust, ashes or other materials. Therefore, a shadow image of a cylindrical sample gets detected at the continuous increasing temperature. Through geometrical changes of the cylinder during the heating, which can be recognized by the projected picture of the sample, it is possible to define the softening point and/or melting point from the considered sample. Such practical results are needed for pyrometallurgical process developments. In the case of desulphurization slag, the hot stage microscope is used to perform lab scale trials. This means, that the cylindrical sample (diameter: 3 mm; height: 3 mm) gets heated (10 K/min.) to defined temperatures and different gas atmospheres. The different continuously added gases, to generate a specific atmosphere in the furnace, were pure oxygen, synthetic air and pure carbon monoxide. To interpret the success of the treatment, the continuously released off-gas composition was analysed and furthermore, a scanning electron microscope analysis from the treated sample was used.

light microscope picture of the treated sample

shadow picture at the end

detected CO2 and SO2 during the treatment

Fig. 2 Main results from the lab scale trials (example for 1400 °C and pure oxygen atmosphere)

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Various trials were performed in lab scale, in the end it was possible to define the temperature range together with the needed treatment conditions for a successful recycling. A representative result for a detailed explanation is shown in Fig. 2. The results were applied in order to define useful process parameter for a successful treatment of the desulphurization slag. This analysis showed, that it is needed to perform such a process in liquid state. This is necessary since lower temperatures (in solid state) lead to a not satisfying sulphur removal. Regarding the detected gases via a gas analyser attached to the hot stage microscope, it was possible to define the best conditions. A representative example is shown in Fig. 2. After a heating up to 1400 °C in an oxidizing atmosphere (provided by the usage of pure oxygen), the carbon from the liquid slag starts to combust and leaves the reaction room with the off-gas stream. After a complete removal of carbon, the sulphur starts to react and forms SO2, which is gaseous as well. The light microscope picture of the treated slag at the mentioned conditions shows a lot of pores, which indicates the gas formation. A subsequent performed chemical analysis of the remaining slags shows sulphur amounts which are lower than 0.01%. These successful investigation and parameter definition in such a small scale leads to an up-scale of the trial size into a technical scale device, which is mentioned in the following chapter.

Recycling Process Development in Technical Scale Size The lab-scale experiments showed, that oxygen in the furnace atmosphere is mandatory to reach a very low level of sulphur in the slag at the end of the treatment. Therefore, a furnace should be used which can supply oxygen with a well mixing effect to homogenize the slag together with the supplied oxygen. Regarding these demands, a TBRC (Top Blown Rotary Converter) is a useful furnace design. Such a TBRC, with a reaction room of about 70 l and an oxygen-methane burner is available at the Chair of Nonferrous Metallurgy, Montanuniversitaet Leoben. Different experiments using the defined process conditions from the lab scale trials were performed in this furnace. The process control was done via a gas analyser in the off-gas stream. Depending on the temperature and CO2 value in the off-gas, oxygen was additional added to the reaction room by a burner which is operated overstoichiometrically condition for a quick and complete sulphur removing. It is possible to control the process with the temperature and the off-gas analyser. The end of the treatment can be easily recognized by the trend of SO2 in the off-gas. This can be shown in the following Fig. 3. A successful sulphur removal from the desulphurization slag can be easily recognized by the characteristic SO2 peak in the off-gas analysis, as it is represented in Fig. 3. After reaching the maximum of SO2, the oxygen is rising in the analysis and indicates that most of the sulphur is removed as well. A range of TBRC trials with different parameter and slag additives to obtain the appropriate viscosity of the slag

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for an ideal treatment were executed. It was possible to confirm the results from the lab scale trial with sulphur levels in the treated slag lower than 0.01% in the technical scale trial with 50–100 kg charged material per batch. To get an impression of the used TBRC in Fig. 4 the tapping after a successful treatment is shown.

Fig. 3 Off-gas analysis from a TBRC trial at 1400 °C

Fig. 4 Tapping of the TBRC after the treatment

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Due to the very satisfying results this process design was verified by pilot scale trial, which was the next step in the development.

Pilot Scale Trial for the Treatment of Desulphurization Slag As next step in this investigation pilot scale trials with 200–1000 kg desulphurization slag were performed in a short drum furnace, which has a similar facility design to the TBRC. To optimize the process itself and also to minimize the treatment time, an additional oxygen lance was installed additionally to the natural gas-oxygen burner. With the individual control of the oxygen flow it is possible to optimize the defined process conditions. Different performed trials in this scale confirmed the results from the former executed investigations. It was possible to verify the feasibility of this process design and the pilot scale trial allowed to define the most useful process parameter.

Conclusion The needed desulphurization of pig iron generates a special type of residue, the so called desulphurization slag. At the moment no useful recycling technology is in operation for this residue, and the desulphurization slag often goes to landfill. To prevent dumping different research projects are ongoing at the moment. In the case of desulphurization slag it was possible to develop an oxidizing recycling process

Fig. 5 Possible process flowsheet for the treatment of desulphurization slag

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within several years starting with fundamental theoretical considerations and lab scale trials, up to pilot scale investigations with 500 kg of treated material. It was possible to remove the sulphur of the slag via the off-gas. The remaining slag will be reused in the primary iron and steel production. A detailed flowsheet about the implementation of this new recycling process is shown in Fig. 5. It is possible to reach sulphur levels in the treated slag which are lower than 0.01% and this can be seen as very good results. This led to a patent which was applied for this process. Due to the changings in the composition and viscosity of the slag different additives were evaluated and the optimum was defined.

References 1. Shi C (1999) Corrosion resistant cement made with steel mill by products. Chinese Society for Metals 2. Damtoft J (2008) Sustainable development and climate change initiatives. Cem Concr Res 115–127 3. Motz H (2000) Products of steel slags an opportunity to save natural resources. Waste Manag 285–293 4. Wegst M, Wegst C (2007) Stahlschlüssel-Taschenbuch, 21st edn. Verlag Stahlschlüssel Wegst, Marbach 5. Gudenau H (2002) Materialsammlung zum Praktikum der Metallurgie, Trans-aix-press, Aachen 6. Hansen M, Elliott RP, Shunk FA (1958) Constitution of binary alloys, 2nd edn. McGraw-Hill, New York 7. Freißmuth A (2004) Die Entschwefelung von Roheisen, 1st edn. Almamet, Ainring 8. Rath Group (2012). http://www.rath-group.com/branchen/fluessigstahl/hochofenauskleidung/ 9. Wiesenberger H, Kircher J (2001) Stand der Technik in der Schwefelsäureerzeugung im Hinblick auf die IPPC-Richtlinie. Umweltbundesamt, Wien 10. Enß J (1990) Aufarbeitung von Entschwefelungsschlacken aus der Stahlindustrie. Chem Ing Tec 62:36–38 11. Schwerdtfeger K, Pawlek F (1983) Berichte der Bunsengesellschaft für physikalische. Chemie 87:1230 12. Reaktoren für heterogene Reaktionen (2012). http://www.chemgapedia.de/vsengine/vlu/ vsc/de/ch/10/heterogene_reaktoren/reaktoren/reaktoren.vlu/Page/vsc/de/ch/10/heterogene_ reaktoren/reaktoren/wirbelschicht_reaktoren/wirbelschicht_reaktoren/wirbelschicht_reaktoren. vscml.html 13. IEHK RWTH Aachen (2012). http://www.iehk.rwth-aachen.de/index.php?id=470 14. Joseph Egli AG (2012). www.jeag.com

High Temperature Phase Formation at the Slag/Refractory Interphase at Ferronickel Production Christoph Sagadin, Stefan Luidold, Christoph Wagner and Alfred Spanring

Abstract Corrosion mechanisms between high melting synthetic ferronickel slags and refractory were investigated. The used slags were prepared by mixing and melting of specific oxides. Substrates of the applied refractory material and specimens of FeNi slags were heated in a hot stage microscope up to 1650 °C. The experiments were performed under a defined gas atmosphere of 60% CO and 40% CO2. A further examination of the formed phases between slag and refractory occurred by scanning electron microscope. The investigations indicate that the slag penetrates between magnesia grains and partly dissolves magnesia. Spot analyses show that iron diffuses into the magnesia grains, which transform to magnesiawustite, meanwhile SiO2 forms different types of olivine like forsterite and monticellite. Thermodynamic calculations confirm the formation of these phases. The combination of practical lab scale experiments and thermodynamic calculations should finally contribute to an improvement of the refractory lifetime and performance.



Keywords Ferronickel Magnesia brick Silica



Refractory



Slag



Corrosion



Ferroalloy

C. Sagadin (✉) ⋅ S. Luidold CD Laboratory for Extractive Metallurgy of Technological Metals, Montanuniversitaet Leoben/Nonferrous Metallurgy, Franz-Josef-Straße 18, 8700 Leoben, Austria e-mail: [email protected] C. Wagner ⋅ A. Spanring RHI AG-Nonferrous Metals, Wienerbergerstraße 9, 1100 Vienna, Austria © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_10

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Introduction Reduced production efficiency and furnace shutdowns are some of the results of inferior refractory materials. In the ferroalloy industry, especially for the ferronickel production, the refractory stability represents an essential factor. The corrosion of refractory caused by chemical attack and high melting slags massively influences the furnace lifetime. Therefore, it is required to reduce the wear of refractory lining to guarantee a safe furnace operation and satisfactory furnace lifetime with short downtimes for repairs [1–3]. The interior brick lining of the electric arc furnace consists of burnt high-purity magnesia refractory bricks, which are commonly used in ferronickel smelting furnaces. Temperatures of more than 1650 °C and high acidic slags are typical for the ferronickel smelting process, which stresses the ceramic magnesia refractory despite its high melting point, good mechanical properties at elevated temperatures and ability to withstand hostile conditions [4, 5]. According to Hu [9], basic magnesia refractory material has a higher resistance performance than acidic refractories in the ferronickel process despite the acidic slags. In detail, FeO and MgO would form a substitution solid solution resulting in a decreased FeO content in the slag system. Consequently, the viscosity of the slag increases and the penetration declines. The current laboratory scale experiment combine practical measurements with theoretical thermodynamic calculations to understand the high-temperature interactions between refractory substrate and synthetically produced ferronickel slag, which is necessary for an improvement of the refractory performance and lifetime [2, 5]. The focus of this work comprises an evaluation of an affected magnesia refractory and the thereby occurring phase formation when a synthetic ferronickel slag interacts with such a substrate. The samples are heated in a hot stage microscope under a defined CO/CO2 mixture to mimic actual process conditions. Further, the examination includes phase determination via scanning electron microscopy (SEM) in combination with thermodynamic calculations by using FactSage 7.1. This approach of practical testing and thermodynamic calculations should ultimately provide a path for improving the refractory lifetime and performance [6–8].

Materials and Methods Hot Stage Microscope The hot stage microscope (HSM) is usually used for the investigation of characteristic temperatures like the deforming-, spherical-, hemisphere- and fluid temperatures from slags, ashes and dusts. The HSM furnace can reach a temperature of 1700 °C with a defined heating rate and investigations can be done under defined

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gas atmospheres. The main components are a halogen lamp, a high-temperature tube furnace, a CCD camera and a control unit. For the investigation, a slag cylinder was pressed out of synthetic slag powder, which was positioned on a refractory plate. With a heating rate of 10 K/min, the high temperature tube furnace was heated to 1700 °C, which correlates to a sample temperature of 1650 °C. Thereby, the slag melts and penetrates the refractory substrate, and react with the high magnesia brick forming different phases. One hour after reaching 1650 °C, the furnace was automatically switched off and cooled down to room temperature. To mimic the ferronickel process in the laboratory-scale test, a gas mixture of 60% CO and 40% CO2 was used for purging the tube furnace at 0.16 l/min. The reducing gas mixture corresponds to an oxygen partial pressure of 1.94 × 10−7 atm at 1650 °C.

SEM/EDS Analysis The refractory substrate including the melted and infiltrated slag was cut by a diamond grinding wheel for cross–section analysis. Then, the sample was embedded in resin (Araldite DBF) and grinded in several steps with different SIC foils (800, 1200, 2400 and 4000). A conductive surface, which is essential for scanning electron microscopy was applied as thin film by sputtering. The slag/ refractory interface microstructure was analysed by a SEM equipped with EDS analyser.

Thermodynamic Investigation The analysis of the interaction between synthetic FeNi slag and magnesia refractory by SEM/EDS were combined with thermodynamic calculations by the software FactSage 7.1. The calculations employed the modules Phase Diagram and Equilib as well as the databases FactPS and FToxide. Focus of the thermodynamic investigations comprised the simulation of refractory attack by slag and its comparison with practical experimental tests.

Materials A. Synthetically produced slag The main oxides in ferronickel slags are SiO2, MgO and Fe2O3 that constitute the oxides for the production of the synthetic slag. 50 g SiO2, 20 g MgO and 30 g Fe2O3 were mixed in a swing mill to get a homogenized oxide mixture. After mixing, the

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Fig. 1 Ternary system of MgO–SiO2–Fe2O3 at 1650 °C in CO/CO2 atmosphere (60/40); the marker represents the slag composition and the dashed line varying mixtures of slag and MgO (as a simplification of the refractory)

oxide mixture was melded in an induction furnace to produce the synthetic slag. Then the slag was ground again in a swing mill to produce a fine slag powder. Figure 1 illustrates the composition of the synthetic slag in the ternary MgO–Fe2O3– SiO2–system at 1650 °C in equilibrium with the CO/CO2 gas mixture, which was compiled by FactSage software. The chemical analysis of the slag is shown in Table 1. B. Refractory substrate A high quality magnesia refractory brick (RADEX S) from RHI Magnesita served as refractory material. Table 1 includes the chemical composition.

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Table 1 Chemical analysis of the synthetic slag and the substrate Chemical analysis

[wt%] SiO2 MgO

CaO

Al2O3

Fe met

Synthetical slag 51.8 19.3 – – 2.7 Refractory brick 0.6 97 1.9 0.1 – (RadexS) BD bulk density (g/cm3), AP apparent porosity (wt%).

Fe2+

Fe3+

Fe2O3

BD

AP

1.8 –

16.3 –

– 0.2

– 3.02

– 15

Corrosion Evaluation by SEM/EDS The corrosion evaluation of the slag/refractory interface by combining HSM and SEM/EDS analyses represents a main part of this research. Figure 2 shows a section of the area near the slag/refractory interface with the slag-melting zone on the top and the refractory material below. The slag infiltrates along the pores up to several mm depth, including the formation of different new phases. In the meanwhile, the corrosion of the refractory starts along the pore, whereas the attack of the slag on magnesia grain remains very low. Figure 3 depicts a magnification of the infiltration (green box) area. It comprises of magnesiowustite and within the smaller pores two olivine phases. Figure 4 shows the element distributions of Mg, Si, Fe and Ca of the slag/refractory

1 mm Fig. 2 SEM/EDS mapping of the cross section area comprising the infiltration and reaction zone between the synthetic slag and the Radex S magnesia refractory brick; Fig. 3 magnifies of the green detail

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Fig. 3 EDS mapping with a line scan

intersection. It shows a qualitative measurement of these elements, wherein brighter-coloured regions correspond to a higher element concentration. The line scan (yellow line in Fig. 3) shows in Fig. 5 the course of concentration over various pores. Table 2 illustrates the content of the oxides in mol% whereby different phases can be determined.

Thermodynamic Calculations In addition to the HSM tests analysed by SEM/EDS, thermodynamic calculations were carried out by the software FactSage 7.1. The model region ranges from pure slag (50% SiO2, 30% Fe2O3 and 20% MgO) to magnesia based refractory (97% MgO, 1.9% CaO, 0.6% SiO2 and 0.1% Al2O3). All thermodynamic calculations applied a reducing atmosphere of 60% CO and 40% CO2 to mimic the practical corrosion test. Figure 6 illustrates the distribution of the elements on individual thermodynamic stable phases as a function of the slag/refractory ratio. The bulk of Si forms with Mg and Fe an olivine phase at 14% refractory in the mixture. At a ratio of 29% refractory material, Fe forms with Mg (refractory) a monoxide phase. The decrease of Fe in the melt causes a simultaneously deceasing of Si, Mg, Ca and Al in the melt up to about 44% refractory. The suppression of Fe and Mg from the olivine causes an intensified enrichment of Ca in the olivine on the basis of the exchangeability of Ca, Fe and Mg ((Mg, Fe, Ca)2SiO4) in this phase. Figure 7 shows for an isotherm section at 1650 °C of the slag/refractory-system the distribution of the components in the stable phases. At low refractory

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Fig. 4 Distribution of Mg, Si, Ca and Fe at the slag/refractory interface analysed by SEM/EDS

Fig. 5 Line scan of the infiltration area, which shows different types of olivine and magnesiowustite; scan 1 = magnesiowustite, scan 2 = monticellite, scan 3 = forsterite

144 Table 2 Results of the linescan

C. Sagadin et al. Element [mol%]

EDS line analyse 1 2

Mg Si Fe Ca (Mg, Fe, Ca)/Si ratio Mg/Fe Mg/Si (Mg + Ca)/Si

48 – 1 – – 48.0 –

15.8 13.1 – 11.6 2.1 – 1.2 2.1

3 27.1 13.2 1.58 – 2.2 17.2 2.1 2.1

Fig. 6 Distribution of the elements (Mg, Fe, Si, Ca and Al) on the stable phases (Melt, Monoxide/ MeO and Olivine) in the range of slag to refractory; calculated at 1650 °C and an atmosphere of 60% CO and 40% CO2

concentration (/(< R> + ) = 0 up to a ratio of 0.15 liquid oxide phase (Slag-liq) and spinel are thermodynamically stable. Afterwards two olivine phases appear besides the liquid phase and spinel in the range from < R>/ (< R> + ) = 0.58 to 0,59. Within the next region from < R>/ (< R> + ) = 0,59 to 0,68 the thermodynamic equilibrium is realized by three solid phases, one spinel and two olivine phases. No liquids are stable. For refractory/slag ratios higher than 0,68, only solid spinel, monoxide and two olivine phases remained at thermodynamic equilibrium.

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Fig. 8 Phase assemblage predicted by thermodynamic simulations for the corroded refractory material as a function of refractory/slag ratio and temperature at p(O2) = 10−6 atm.: Slag-liq— liquid oxide, Monoxide solution, Spinel and Olivine

Validation with Post Mortem Analysis The microstructure of a post mortem sample taken from a Peirce Smith converter was examined to understand and corrosion mechanisms and validate the simulations. It has to be mentioned that for the thermodynamic calculation no contact with matte was considered, thus no sulfur attack of the brick was simulated. Due to the batch processing and different process steps of iron blowing and copper blowing the simulation and validation with post mortem samples is generally difficult. Furthermore, it has to be noted that conducted EDS analysis are not obtained from a single spot but from an interacting volume in the μm scale. Because of the poor spatial resolution of the analyses the spot analyses are not corresponding to expected ranges of stoichiometry/chemistry for known crystalline phases. Thus, a comparison of the element distribution of the spots did not give an exact quantitative analysis but good qualitative statements of the differences between spots in the area studied. The interface between refractory material and molten fayalite slag showed a slag coating on the hot face and infiltration, which can be considered as the reaction layer, Fig. 9. Several horizontal cracks partly filled with copper and matte are

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Fig. 9 Cross section post mortem sample and phase distribution from hot to cold face, calculation based on SEM-EDX area analyses [x-axis in mm from hot face towards cold face]

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visible. Joints are also filled with metallic copper. Additionally an analysis of scans distributed over the infiltrated area was performed to identify the phases taking place at the expected corrosion reaction of the refractory substrate by fayalite slag.

Results and Discussion Thermodynamic equilibrium calculations combined with post mortem analysis, were applied to investigate corrosion mechanisms between refractory materials and molten slags used in copper smelting. Thermodynamic calculations with varying temperatures and refractory/slag ratio were conducted. The characteristic of the phases predicted by the thermodynamic calculations corresponded with the assumptions and post mortem results. It was shown that the results from fayalite slag corroding the refractory substrate as well as the thermodynamic calculations were in good agreement concerning formed phases and their compositions. The main corrosion mechanism was wear by spalling of degenerated brick parts. Several cracks partly filled with slag and matte were observed. In addition the brick showed chemical attack due to high SiO2 and sulfur supply resulting in the formation of forsterite and monticellite. Although physical parameters such as the granulometric differences between the aggregates and the matrix, the location of these phases in the refractory microstructure and the open porosity and permeability of the refractory material were neglected, the thermodynamic calculations reflect the results from post mortem analysis and vice versa. Therefore, the thermodynamic calculations are suitable for predicting the sequence of phase transformations in equilibrium and therefore show the usefulness of this tool for supporting experimental work. By combining the results of post mortem investigations and the thermodynamic analysis using FactSageTM, refractory corrosion mechanisms were investigated to draw implications for improving the refractory performance and lifetime by means of changing physical properties or chemical compositions of the bricks. Ongoing work focuses on different refractory materials, slag systems as well as thermodynamic calculations investigating refractory corrosion mechanisms.

Conclusion A thermo-mechanical and a thermo-chemical simulation case study were discussed. Both cases show the importance of numerical simulations and the applicability for solving refractory design related problems. Simplified model approaches were used to obtain results in reasonable time scales and to reduce modelling efforts.

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Due to the complexity of refractory corrosion the validation of models show rather a qualitative than a quantitative agreement. The discussed thermo-mechanical simulations make use of a linear elastic material model and do not consider the change of the material properties due to infiltration of pores and chemical reactions. Those facts typically lead to overestimated stress magnitudes, but the general fracture pattern of a brick complies with the calculated tensile stress fields. Thermo-chemical calculations predict the phase and liquid slag formation due to chemical reactions between slag and refractory. They are used to show the qualification of a refractory material for a specific process. The calculation results show a fairly good agreement with the phases found in a post mortem sample. As the standard approach shows no spatial discretisation it cannot be predicted where the phases are formed and which refractory/slag ratio actually exists. Furthermore, the simplification of an established thermodynamic equilibrium might lead to differences between the real corrosion conditions and the simulation results Development of advanced industrially applicable simulation models is needed improve the refractory engineering process as simulations today are more of indicative nature and might lack in accuracy. Currently following developments are promoted for enhanced thermo-mechanical and thermo-chemical simulation capabilities in the field of refractory design: • Development of solidification models for metal and slag • Improving the physical property database for refractory materials • Implementation, verification and validation of refractory models considering the non-linear strain, temperature and time dependent behaviour of refractory material • Introduction of smeared material models which behave phenomenological like a brick lining but neglecting the high number of contacts for simulation of linings with several hundred to over thousand bricks • Developing pore models for thermo-chemical corrosion simulations for a spatial discretisation • Considering the kinetics of refractory corrosion.

References 1. Schlesinger ME, King MJ, Sole KC, Davenport WG (2011) Extractive metallury of copper. Elsevier Ltd, Oxford 2. Wang S, Davenport W, Siegmund A, Yao S, Gonzales T, Walters G, George D (2016) In: 9th International copper conference copper smelting: 2016 world copper smelter data, Kobe, 13–16 Nov 2016 3. Routschka G (2017) Praxishandbuch Feuerfeste Werkstoffe. Vulkan-Verlag, Essen 4. Wagner C, Wenzl C, Gregurek D, Kreuzer D, Luidold S, Schnideritsch H (2017) Thermodynamic and experimental investigations of high-temperature refractory corrosion by molten slags. Metall Mater Trans B 48B:119–131

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5. Lee WE, Zhang S (1999) Int Mater Rev 44(3):77–104 6. Bale CW, Chartrand P, Degterov SA, Eriksson G, Hack K, Ben Mahfoud R, Melancon J, Peltron AD, Peterson S (2002) CALPHAD 26(2):189–228 7. Berjonneau J, Pringent P, Poirier J (2009) Ceram Int 35(2):623–635

Wear Phenomena in Non-ferrous Metal Furnaces D. Gregurek, C. Majcenovic, K. Budna, J. Schmidl and A. Spanring

Abstract In non-ferrous metal furnaces the installed magnesia-chromite refractory lining is exposed to several stresses, rather complex in their interaction. Therefore, a detailed investigation and understanding of wear mechanisms through “post mortem studies” is an important prerequisite for refractory producer. Additionally, in order to determine the most suitable refractory products and to improve the lining life of refractories, practical corrosion testing with processing slags is performed. For this purpose the test-facilities, such as induction furnace, rotary kiln but also cup test and drip slag test, allow the best possible understanding of chemothermal brick wear on pilot scale. Prior to testing a complete mineralogical investigation, thermo-chemical calculation via FactSageTM of the slag was carried out. Based on such research results, combined with specific process knowledge RHI Magnesita can recommend appropriate brick lining solutions for non-ferrous metal furnaces. Keywords Refractories



Wear phenomena



Experimental testing

Introduction The profitable pyrometallurgical operation depends on many factors such as furnace type, process conditions, lining design, selection of refractory types, etc. In the non-ferrous metals industry, particularly in copper and lead smelting furnaces, magnesia-chromite bricks are the preferred refractory choice due to their high corrosion resistance [1]. Nevertheless, the refractory lining is exposed to complex and mutual wear caused by chemical, thermal, and mechanical stresses [2]. Therefore the exact understanding of the wear phenomena through post mortem studies is an important prerequisite for the refractory producer, since it provides the D. Gregurek (✉) ⋅ C. Majcenovic ⋅ K. Budna RHI Magnesita, Technology Center Leoben, Magnesitstrasse 2, Leoben 8700, Austria e-mail: [email protected] J. Schmidl ⋅ A. Spanring RHI Magnesita, Wienerbergstrasse 9, Vienna 1100, Austria © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_13

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basis for both customer recommendations and innovative product development. In addition to post mortem studies, laboratory work and experimental testing carried out in the pilot plant of RHI Magnesita Technology Center Leoben enable the best possible understanding of the brick wear on the pilot scale [3]. Combining all these facts, the adequate choice of refractory is always essential for a successful furnace campaign. A general overview of wear phenomena in the non-ferrous industry was discussed and introduced in several papers in the past [4–7]. Malfliet et al. carried out a critical review work on degradation mechanism and use of the refractory linings in the copper production processes [4]. These findings for copper slags are also interesting and relevant for a better understanding of the refractory wear processes in lead metallurgy, as similar slag systems and refractory qualities are used. Particularly in the lead industry a lot of work was done regarding refractory corrosion testing in different pilot-scale and industrial furnaces. For instance, Oprea discussed failure mechanisms observed on the magnesia-chromite bricks lined above the slag line of the flash furnace for zinc-lead smelting [5]. In order to explain these findings some laboratory work was done additionally. Prestes et al. analyzed wear phenomena on magnesia-chromite bricks from the lead short rotary furnace [6]. Similar to Oprea, in addition to post mortem studies also some experimental work by crucible corrosion testing was carried out. Wei reviewed available literature concerning corrosion of refractories in the lead-smelting reactors such as KIVCET furnace and TBRC and evaluated in the laboratory the corrosion behaviour of various refractory materials against industrial slags [7]. The present work gives and overview of the most common refractory wear as observed in non-ferrous production furnaces, using samples from different production stages and metallurgical furnace types. Additionally the paper describes the recent test work carried out by RHI Magnesita on refractory bricks and customer provided slags.

Analytical Procedure Every single post mortem study starts with the visual inspection carried out on the brick cut section followed by selection of samples for chemical analyses and mineralogical investigation. The chemical analyses are carried out by using X-ray fluorescence analysis (Bruker S8 TIGER). The mineralogical investigation is performed on polished sections using reflected light microscope, X-ray diffraction (Bruker D8 ADVANCE), and scanning electron microscope (SEM) (JEOL JSM-6460) combined with an energy-dispersive and wavelength-dispersive X-ray analyser. The measurement of the melting point of slag using a heating stage microscope was carried out according to DIN 51730. This information is particularly important in subsequent testing of slag/refractory interactions, namely to correctly adjust the temperature in the testing furnace.

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Refractory Wear In the present work the most common wear phenomenon occurring in the non-ferrous furnaces such as slag attack, sulphur corrosion and metal infiltration will be discussed in detail.

Chemical Slag Attack The most frequent wear of the refractory lining in the copper and lead furnaces is corrosion by acidic slag. The chemical composition of the main slag types from non-ferrous metallurgical vessels can be roughly expressed by the system FeO-CaO-SiO2 (Fig. 1). Generally the corrosion of the refractories by slag attack manifests itself in three ways [9]: (a) Dissolution reaction occurring at the immediate brick hot face: The driving force here is the lower activity of the refractory oxides like MgO in the slag. The dissolution process, at least in the closed system, will continue until the liquid slag has reached saturation. However, in practice, the point of saturation is never reached and dissolution continues until the entire refractory has been consumed. (b) Dissolution and chemical reaction within the refractory microstructure: Infiltrating slag will dissolve magnesia especially from the fines according to the respective phase equilibrium. This will not directly contribute to corrosive wear which takes place at the immediate refractory hot face. Nevertheless it will contribute to wear by preparing hot erosion due to a loss of brick bonding. (c) Kinetics of slag infiltration: Kinetics of slag infiltration causing processes mentioned above depends on several parameters like viscosity, pore size distribution and wetting angle.

Fig. 1 Typical chemical compositional range of lead- (left side) and copper slags (right side) is demonstrated in CaO-FeO-SiO2 phase diagrams [8]

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At a microscopic level, several zones can be distinguished, as shown in Fig. 2. Below the slag coating at the immediate brick hot face, a thin reaction zone is usually followed by an infiltrated and corroded brick microstructure. Within the reaction zone, the magnesia brick component is frequently dissolved leaving relics of chromite as well as primary and secondary chromite precipitations. Below the reaction zone in the infiltrated and corroded brick microstructure, due to corrosion of the brick-inherent magnesia (coarse grains and matrix fines), the main reaction products include (calcium)-magnesium silicates such as forsterite (Mg2SiO4) and monticellite (CaMgSiO4). The interstitial phase of the magnesia component, especially the di-calcium silicate (Ca2SiO4), is also corroded. Generally the periclase (MgO) is more basic than chromite and therefore more susceptible to acidic corrosion. The chromite is not corroded but modified in chemical composition due to diffusion phenomena (usually enriched with iron-, zinc- and copper-oxide).

Fig. 2 Microphotograph of the immediate brick hot face taken with reflected light microscopy. Microstructural overview of a used magnesia-chromite brick from a flash smelting furnace. Slag coating (S). Reaction zone (R). Infiltrated and corroded brick microstructure (I)

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Sulfur Attack Another very common type of chemical attack is corrosion by sulfates usually in the gas zone. The penetration of gaseous SO2 caused by oxidation of sulfidic ore leads to the formation of SO3, which below temperatures of approximately 1050 °C reacts with the basic oxide of the magnesia-chromite brick leading to the formation of magnesium sulfates (MgSO4) (Fig. 3). The fundamental reactions are shown below [2], MeS + O2 → MeO + SO2 MeS + MeO → Me + SO2 SO2 + 1 ̸ 2 O2 ↔ SO3 ð ∼ 760 ◦ CÞ SO3 + MgO → MgSO4 ð < 1050 ◦ CÞ SO3 + CaO → CaSO4 where MeS corresponds to metal bearing sulphides like Cu2S, PbS, etc. Although the oxidation of SO2 to SO3 rapidly decreases above 760 °C, a certain partial pressure of SO3 can be assumed in the temperature range between 760 °C and 1050 °C that would allow the formation of basic sulfates. The intensity of sulfate corrosion generally depends on many factors such as the amount of supplied SO2, the surplus of acidic SO2 versus the basic components of

Fig. 3 Microphotograph taken by scanning electron microscopy showing the microstructural details of a used magnesia-chromite brick from a Peirce-Smith converter. Corroded magnesia (1). Chromite (2). (Ca)-Mg-sulfate (3). Pore (4)

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the infiltrate (e.g., alkalis and CaO), reaction temperature and time, as well as brick properties such as porosity and brick composition (e.g. lime to silica ratio, CaO/ SiO2). In case of a magnesia-chromite bricks with high CaO/SiO2 ratio also the interstitial phase such as di-calcium silicate is massively corroded. The latter results in formation of Ca-sulfate (CaSO4). Due to severe corrosion of the brick bonding phase the initially high CaO/SiO2 ratio is decreased into the stability area of forsterite. Therefore numerous single forsterite crystals could additionally form in such melt. This means that the CaO/ SiO2 ratio of the silicate phases is shifted to a lower value, and this similarly proceeds until forsterite is formed. At temperatures above 1050 °C, MgSO4 will dissociate and form fine crystalline MgO at magnesia rims. This takes place without rebuilding of the original ceramic bonding of the brick.

Non-oxide Infiltration In addition to acidic slag also other components like metallic copper, lead, matte, etc. infiltrate the brick microstructure (Fig. 4). Similar to the acidic slag, the degree of infiltration depends on the surface tension, the boundary angle in contact with the refractory oxides, the metal density, the bath height and the size distribution of the brick pores [2].

Fig. 4 Microphotograph taken by reflected light microscopy showing the brick microstructure of a used magnesia-chromite brick from a smelting furnace, which is completely infiltrated with nickel-copper-cobalt matte. Magnesia (1). Chromite (2)

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Experimental Testing In addition to the experience acquired through post mortem studies described in previous chapters, practical knowledge from experimental trials is necessary for the appropriate refractory product selection. To achieve this goal, test procedures have been developed and installed at the RHI Magnesita Technology Center in Leoben, Austria, to simulate the various demanding service conditions [3]. Throughout a co-operation projects with RHI Magnesita customers practical tests were carried out with different refractory types—usually pre-selected magnesia-chromite and alumina-chromia bricks. For example static and dynamic corrosion test methods and equipment such as the cup test, induction furnace, rotary kiln, drip slag test, high frequency induction furnace enable a comprehensive understanding of brick wear following such pilot scale trials. Exemplary testing work carried out in the induction furnace is shown in Fig. 5.

Slag Characterization Prior to experimental testing, the initial characterization of slag including complete chemical and mineralogical analysis, followed by a determination of melting points measured with a heating microscope and thermochemical calculations with FactSageTM software [10] was performed. The chemical analysis of slag is shown in Table 1. An example of a slag characterization received from the Peirce Smith converter is described below. The main component in crystalline slag was fayalite (Fe2SiO4) enriched with zinc. Additionally, slag contains Zn-Sn-Ni-Al–Fe-oxide of magnetite type (Fe3O4), minor amounts of cristobalite (SiO2), and an amorphous glassy phase. Generally, the form of Cu-bearing minerals depends on the copper slag type and the basic process route in combination with its specific raw materials (i.e., primary or secondary metal production) [11]. They can be in the form of oxides or sulfides or a mixture of both. In the present slag the main form of copper was Cu-sulfide (Cu2S). Determination of the melting point by heating microscope revealed that the slag was completely liquid at approximately 1420 °C in air. The thermodynamic calculations were performed with the FactSageTM 7.1 software by using FactPS, FToxide and FTmisc databases. The calculated liquidus temperature at Po2 = 10−6 atm is 1350 °C (97% molten slag; 3% matte, Fig. 6). The solidus temperature corresponds to ∼800 °C (at Po2 = 10−6 atm).

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Fig. 5 Photographs showing corrosion test work carried out in the induction furnace at the RHI Magnesita Technology Center Leoben

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Table 1 Chemical slag analysis (wt%) MgO

Al2O3

SiO2

CaO

Fe2Oa3

Na2O

1 4 30 2 45 2 a total iron and copper calculates as Fe2O3 and CuO

SO3

CuOa

ZnO

SnO2

PbO

3

5

6

1

2

100 90

ProporƟon of phases [wt%]

80 70 60 50 40 30 20 10 0 800

900

1000

1100

1200

1300

1400

1500

Temperature [°C] MaƩe

Slag

Spinel**

Corundum**

Solid*

Fig. 6 FactSageTM calculation of the slag liquidus and solidus temperatures performed at pO2 = 10−6. Solid* means solid phases such as SiO2, Na-K-Al-silicate (feldspar, NaAlSi3O8) and Cu-Fe-oxide; **Other solid phase include also corundum and spinel

Induction Furnace Testing The refractory wear during induction furnace testing is principally dependent on the chemical composition of the customer slag, the chemical composition of the refractory, testing temperature, and trial duration. In order to generate maximal refractory wear in a relatively short testing time, the furnace operating temperature is usually higher than that occurring during the actual customer process. The induction furnace can be lined with up to 16 bricks (4–5 different brick brands). The furnace diameter is 250 mm and the temperature can range between

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1500 °C and 1750 °C. Up to 60 kg of metal and 1.5–3 kg of slag (per cycle) were used during the test. The process slag was changed every 0.5–1.5 h to achieve a constant, highly corrosive medium in contact with the installed refractory material. Testing is usually performed for 6–8 h. Figure 7 shows an example of a refractory wear after a test performed in the induction furnace with a fayalite slag type at a temperature of 1650 °C. The refractory is partially covered with a thin slag coating. The highest macroscopically visible wear is always at the interface between the refractory, metal, and slag which can be explained by the Marangoni convection. The latter phenomenon occurs at the phase boundaries and is driven by a gradient of the interface tension of the fluid phases [9]. In metallurgical cases normally the three phase boundary metal/slag/refractory is causing a more severe convection than the boundary slag/gas phase/refractory [12]. An example of the microstructural investigation after testing in the induction furnace is shown in Fig. 8. The brick microstructure at the immediate refractory hot face is covered with thin slag coating respectively reaction zone. Below that area an infiltrated corroded brick microstructure could be observed. In the infiltrated brick microstructure, corrosion of the magnesia component took place (Fig. 8). The main reaction product is forsterite.

Fig. 7 Photograph showing cut section of a magnesia-chromite brick after an induction furnace test. The immediate brick hot face is partially covered with slag. Sample for mineralogical investigation (1) (see also Fig. 8)

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Fig. 8 Microphotograph of the infiltrated magnesia-chromite brick after the induction furnace test taken with reflected light microscopy. Corroded magnesia component (1). Chromite (2). Forsterite (3)

Conclusions The wear mechanisms discussed in this paper demonstrate that a combination of slag attack, infiltration and corrosion of the bricks inherent components lead to a softening of the bricks microstructure, loss of flexibility and brick strength. This weakened microstructure is then susceptible to continuous wear by hot erosion. Additionally, due to the changes in the thermo-mechanical properties of the refractory, thermal shock leads to crack formation, primarily at the interface between the infiltrated and non-infiltrated brick areas, and finally, to discontinuous wear by spalling. The high sulfur supply, typically occurring when processing sulfidic materials under oxidizing conditions, leads to corrosion of the brick-inherent magnesia and of the interstitial CaO-containing secondary phase within the magnesia. Chromite generally shows a higher corrosion resistance against both fayalite slag and sulfur attack. Non-oxide infiltration of the brick microstructure by metals or sulfidic components (matte) dramatically changes the thermal conductivity of the brick, thus increasing the susceptibility to crack formation and spalling, which is intensified by thermal shocks. A combination of different experimental testing methods enables the best possible understanding of chemical brick wear on a pilot scale. In the present paper example of the corrosion testing work carried out in the induction furnace was described in detail. In order to generate maximal refractory wear in a relatively short testing time, the operating furnace temperature was higher than that at the customer

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site. The initial characterization of slag including chemical and mineralogical analysis, determination of melting points and thermochemical calculations are important prerequisite prior to any experimental testing. A detailed investigation of wear mechanisms in combination with results of experimental testing are a vital tool for the refractory producer. They provide the basis for both product recommendations and innovative product development. The post mortem investigations can clearly highlight which specific stresses affect the refractory products in the various non-ferrous metal processing furnaces. Based on the investigation results, in combination with the long-term service experience, a refractory producer can recommend the most appropriate choice of furnace lining. This is frequently enhanced through active collaborations with the customer.

References 1. Routschka G (2017) Praxishandbuch Feuerfeste Werkstoffe. Vulkan-Verlag, Essen 2. Barthel H (1981) Wear of chrome magnesite bricks in copper smelting furnaces. Interceram 30:250–255 3. Gregurek D, Ressler A, Reiter V, Franzkowiak A, Spanring A, Prietl T (2013) Refractory wear mechanism in the nonferrous metal industry: testing and modeling results. JOM 65 (11):1622–1630 4. Malfliet A, Lotfian S, Scheunis L, Petkov V, Pandelaers L. Jones PT, Blanpain B (2014) Degradation mechanism an use of refractory linings in copper production processes: a critical review. J Eur Ceram Soc 34:849–876 5. Oprea G (2004) Failure mechanism of refractory linings for non-ferrous flash smelting furnaces. In: Proceedings of tehran international conference on refractories, 4–6 May 2004 , Tehran 6. Prestes E, Chinelatto ASA, Resende WS (2009) Post mortem analysis of burned magnesia-chromite brick used in short rotary furnace of secondary lead smelting. Ceramica 55:61–66 7. Wei L (2000) Corrosion of refractories in lead smelting reactors. MSc thesis, University of British Columbia. 8. Osborn EF, Muan A (1960) Phase equilibrium diagrams of oxide systems. American Ceramic Society with the Edward Orton Jr. Ceramic Foundation, Columbus, Ohio 9. Harmuth H, Vollmann S (2014) Refractory corrosion by dissolution in slags—challenges and trends of present fundamental research. Iron Steel Rev 58(4):157–170 10. Bale CW, Chartrand P, Decterov SA, Eriksson G, Hack K, Mahfoud BR, Melançon J, Pelton AD, Petersen S (2002) FactSage thermochemical software and databases. Calphad J 62:189–228 11. Davenport WG, King M, Schleisinger M, Biswas AK (2002) Extractive metallurgy of copper. Kidlington, Elsevier Science Ltd, Oxford OX5 1 GB, UK 12. Pötschke J, Brüggmann Ch (2012) Premature wear of refractories due to marangoni-Convection. Steel Res Int 83(7):637–644

A Scientific Roadmap for Refractory Corrosion Testwork J. Schmidl, A. Spanring, D. Gregurek and K. Reinharter

Abstract In non-ferrous metallurgy the service life of refractory materials typically ranges from several weeks up to three years or more and is strongly dependent on operating conditions. To support our customers with the most viable refractory solution for their needs RHI Magnesita follows a structured approach of thermo-chemical calculations, experimental evaluation from lab scale up to industrial field tests and post-mortem analysis of used refractory materials. This paper will give an overview of the RHI Magnesita refractory selection process and the most recent developments of the experimental setup for the so called HF-IF test. This is a dynamic corrosion test that allows to determine fundamental properties of different refractory material types under process conditions subjected with respective customers process materials. The HF-IF experimental procedure is an excellent tool to simulate refractory wear in industrial processes, diminishing risks associated with plant trials and support decision making to choose the optimal refractory solution for the customer. Keywords Refractory



Corrosion testwork



HF-ITO test

Introduction RHI Magnesita provides refractory solutions to almost all metallurgical processes in the nonferrous metals industry. An average sales volume of more than 90.000 t/a allows RHI Magnesita to serve customers of the non-ferrous base metals (Cu, Pb, Zn), ferro-alloys (FeCr, FeNi, FeMn, …) and precious metals industry [1]. Both, primary and secondary smelting operations are facing the challenge to treat especially complex material feed mixes, often at varying feed mix compositions. This J. Schmidl (✉) ⋅ A. Spanring RHI Magnesita, Wienerbergstrasse 9, Vienna 1100, Austria e-mail: [email protected] D. Gregurek ⋅ K. Reinharter RHI Magnesita, TCL, Magnesitstrasse 2, Leoben 8700, Austria © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_14

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changes of input parameters influences not only the interaction with the refractory lining but also the output of the smelting and/or recycling process. Existing results of a process cannot be directly transferred to describe other processes with different input parameters although they might be similar at a first glance. This means that every request needs to be investigated to provide customers with the best available solution. Rather than focusing on trial and error RHI Magnesita follows a well-structured scientific approach that combines metallurgical and refractory expertise to select the best suitable refractory material for a given application. Typically stresses in the refractory lining can be subdivided into chemical, thermal and mechanical stresses which will affect the service life of the furnace lining and have been published for a number of different applications in literature [2–7]. Therefore several corrosion testing methods such as the induction furnace, short rotary kiln, cup test, etc. have been developed to assess different refractory brand qualities in order to recommend industrial trials at customer plants, leading finally to an improved refractory performance. The latest development in this series of high-temperature testing methods is the so called HF-ITO test which capabilities will be described together with the generic approach for refractory selection in this article.

Problem Formulation At the very beginning the crucial part in the selection process of a refractory brand quality for an application request is the problem formulation. This part is the specific distinction between “collect all the information you can about a subject” and “let’s do an experiment”. Due to the large number of input variables and the nature of their relationship, research problems in the field of slag|matte|metal interaction with refractory materials are rather complex. To understand the dimensions of a problem it is necessary to consider focus groups of variables to get an insight into a specific set of questions from the process. Apart from obvious information like brand quality and service life history it is therefore particularly important to have information about the slag|matte|metal|offgas composition, process temperature and their variations in composition during the expected life time of the refractory lining. The collected information is then used to estimate the phase equilibria using the actual slag|matte|metal composition from industrial samples at process temperature and calculate the oxygen (pO2) and sulphur (pS2) partial pressures of the system. Figure 1 shows the sulphur-oxygen potential diagram for copper production [8]. This diagram can be understood as kind of a navigation map indicating the common working region that has to be considered during the copper production process, starting from smelting over converting, fire refining up to slag cleaning, Cu-recycling or smelting of complex CuPb containing mattes [9]. It should be emphasised that other impurities will impact the working regions and therefore

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Cu-recycling fire refining

PSC - copper blow

PSC - slag blow maƩe smelƟng

Cu slag cleaning slag fuming

CuPb maƩe smelƟng (blast furnace, EAF)

Fig. 1 Sulfur-oxygen potential diagram for the system Cu–Fe–S–O–SiO2 at 1250 °C

thermodynamic equilibrium calculations are used to predict the actual pO2/pS2 combination for the process to use this information during corrosion test-work.

Post-mortem Studies Understanding the refractory wear phenomena through post-mortem studies on used refractory bricks is an essential tool as they provide a precise understanding of the wear parameters influencing the refractory performance on the one hand and provide impetus towards the development of new refractory products. During relining the post-mortem samples are taken, quite frequently supported by RHI Magnesita experts, and their position in the furnace is documented. Together with the process information this allows a precise description of the wear

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mechanisms of the spent refractory sample. From the macroscopic appearance typically four zones can be observed • • • •

accretion layer of slag|matte|metal from the process reaction layer infiltration zone original microstructure of the brick up to cold side

The bricks are then sliced and cut into smaller samples to be prepared for further microscopic examination and analysis of the chemical composition. The basis for the evaluation of the refractory wear is an infiltration diagram (Fig. 2). It shows the mean phase distribution as a function of the distance from the hot face depending on the thermal gradient, thermal profile and is based on chemical and mineralogical, especially microscopic investigation. The diagrams are designed to follow the evolution of phase distribution starting from the hot face up to the cold face and can be discussed to determine following aspects of refractory degradation • liquid infiltration by slag

phase fracƟon in %

While significant content of iron oxide is found together with relatively less silicates this may mean that the slag is not especially aggressive towards magnesiachromite bricks, since free silica would probably react with the excess iron oxide. The low depth of infiltrated slag also shows that slag may play only a secondary role on the wear of the bricks in the converter.

distance from hot side in mm

Fig. 2 Phase distribution of a post-mortem refractory sample from hot to cold face, calculation based on SEM-EDS area analyses

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• liquid infiltration by matte and copper It is also noticeable that metallic copper infiltrates deeper into the brick than copper matte: this can be the result of the slightly lower melting temperature of copper (1085 °C) in comparison to Cu2S (1130 °C). Another observation is the presence of copper oxide (Cu2O) at the end of the infiltrated zone coexisting with metallic copper. Oxides of copper are not observed in the hottest parts of the brick. Because its it is clearly not coming from the slag, the most likely conclusion is that the copper oxide is being formed in situ. Associated to the fact that metallic copper content increases when matte decreases and infiltrates deeper, it is hereby postulated that matte infiltrates in larger amounts than copper, and that most of the copper present in the structure of the brick (not in the cracks or joints) is formed by oxidation of matte according to following reactions Cu2 S + 1.5 O2 = 2 Cu + SO3 2Cu + 0.5O2 = Cu2 O

ΔV = − 48%

ΔV = + 62%

Cu2 S + 2O2 = Cu2 O + SO3

ΔV = − 15%

ð1Þ ð2Þ ð3Þ

If only the direct oxidation of copper sulfide (density 5.8 g/cm3) to copper oxide (density 6.15 g/cm3) is observed according to (Eq. 3) it should result in no loss of structural integrity. However, if reaction occurs according to the two first steps, whereby there is a large shrinkage caused by formation of metallic copper (density 8.9 g/cm3), followed by a large volume increase as a result of its oxidation. Theoretically, this volume increase could be accommodated in the voids previously filled by the infiltrated matte (the net volume increase is given by the third equation, meaning that the oxidation of copper generated by oxidation of matte will occupy less volume than the original matte). However, as copper is formed, voids are open to the infiltration of more matte or metal and the formation of Cu2O may increase damage by spalling. • redox reaction by sulphate diffusion The microstructural studies have confirmed the presence of three sulfates from CaO and MgO. The melting temperature decrease in the order of CaSO4 (TM > 1400 °C), CaMg3(SO4)4 (∼1200 °C) and MgSO4 (∼1135 °C). It is clear from the analysis, that CaSO4 is present in the hottest parts of the bricks and MgSO4 in the coldest. The double sulfate appears, when present, between both sulfates or up to the cold face. This distribution suggests that sulfates might be good tracers to determine the thermal gradient inside the lining.

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Refractory Quality Pre-selection The appropriate tool to determine the chemical suitability of a refractory quality for a given application is a thermochemical calculation using for example thermochemical computational tools such as FactSageTM [10]. This computation method indicates which phases reach thermodynamic stability and, therefore, can be formed by the interactions between refractory and slag. The calculations are performed at isothermal conditions and fixed oxygen partial pressure. A typical representation of the result is shown in Fig. 3. Starting from a fully liquid slag at the given process temperature and oxygen potential, refractory material is added stepwise and the thermodynamic equilibrium is calculated for each single step. The grey line indicates the ideal behaviour, assuming that no interaction between slag and refractory takes place. The results are represented by illustrating the formation of liquid slag and solid phases at equilibrium condition. For the illustrated system three phases are stable at equilibrium – liquid slag, spinel and monoxide. Dissolution of the refractory material into the slag phase is indicated by an increase of the liquid slag area above the grey line. This mechanism is called active corrosion or direct dissolution in literature. If the refractory|slag system tends to form solid precipitates the liquid slag area at equilibrium is beneath the grey line which is called passive corrosion or indirect dissolution. While it is obvious that active corrosion will directly force the refractory degradation this cannot be easily derived from equilibrium calculations for passive corrosion. The formation of new phases is accompanied with volume changes and depending on the actual quantity of this change and the position inside the refractory brick resulting from infiltration of liquid slag into pores. This corrosion

Fig. 3 Typical representation of thermochemical equilibrium calculations in the slag|refractory system carried out with FactSageTM

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mechanism could also lead to mechanical degradation by forming cracks and ultimately spalling off of larger brick parts.

Corrosion Testwork A variety of different corrosion test methods are available at RHI Magnesita and selected according to the findings of the above mentioned results. Special emphasis shall be given to the so called HF-ITO method. This test is especially designed to generate maximal refractory wear in a relatively short testing time in contrast to pilot scale tests in an induction or a short rotary furnace [11] and enables a good pre-selection of suitable brand qualities. Based on this evaluation a decision can be made to either perform pilot scale tests at the RHI Magnesita Technology Centre or directly run a field trial at the industrial application. The degradation of the refractory in contact with liquid slag can be expressed by following equation Ji =

D ⋅ Δci δ

ð4Þ

where Ji is the dissolution rate of the refractory material which is assumed to be controlled by diffusion and D is the diffusion coefficient, δ is the diffusion boundary layer thickness and Δci represents the concentration gradient of the actual refractory component in the slag versus the saturation solubility in the slag. While thermodynamic calculations only represent the phase equilibrium of the slag|refractory interface which accounts for Δci, this test method enables • to simulate agitation of the liquid bath, influencing the diffusion boundary layer • to assess differences between types of raw materials of the refractory type (sintered, fused, pure/mixed, OXICROM, …) by macro- and microscopic examination of the samples after testing • compare the performance of up to four different brick qualities under standardized conditions Both parameter are essential information in the selection process and the HF-ITO method is not only a well-suited scientific test but also an economically solution as it can deliver insights and results in a short period of time and thus resulting in shorter response time to the customers. The original test set-up is illustrated in Fig. 4. The principle test set-up was installed in December 2013 as a stand-alone device, equipped with an offgas unit. The major unit is an induction furnace with a power supply of 40 KW that can be run at a frequency of 5–100 kHz. The furnace can operate at a maximum process temperature of 1700 °C and a rotation speed of the sample holder from 1 to 10 rpm. Up to 3 kg of test material, industrial or synthetic slag|matte|metal samples can be melted in the crucible. The standard crucible material is graphite but, depending on sample composition other crucible materials

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Fig. 4 Original HF-ITO test set-up

have to be considered. The sample holder is placed above the crucible to pre-heat the sample finger and then immersed into the liquid bath. Four samples can be investigated in parallel, each of it with a dimension on 20 × 25 × 115 mm. The temperature measurement is taken from the outer sidewall of the crucible using a Metis MQ11 pyrometer. Compromises have to be made to the furnace atmosphere. It is not possible to accurately control the oxygen partial pressure, two options are available—the test can either be performed in air or inert atmosphere using Argon as purging gas. The testing period typically is around 4 h but can be adapted according to the need of the experiment. To improve significance of the obtained results from the HF-ITO test further developments in both the operation procedure and the test set-up are in progress. Figure 5 shows a sketch of the current HF-ITO test design. • crucible material—the selection of the crucible material is of high importance. When using standard graphite crucibles, interactions of the test material change its composition. Due the small melting volume this could have a high impact on the obtained results. Therefore crucibles are lined with refractory material that ensures minimum interaction with the test material. • atmosphere control—in order to be able to control and adjust a certain pO2/pS2 potential the furnace atmosphere can be purged with pre-mixed gas mixtures. • Temperature measurement—installation of a thermocouple directly in the melt for precise measurement of the actual temperature • sample quenching—after designated time the samples are pulled out of the bath and quenched as fast as possible in order to represent equilibrium/process conditions in the final sample for microscopic investigation.

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Fig. 5 Sketch of the current HF-ITO test design

Conclusion RHI Magnesita follows a well-structured scientific approach that combines metallurgical and refractory expertise to select a refractory material for a given application. This includes detailed know-how of the metallurgical process conditions, collecting information from post-mortem analysis, thermochemical considerations, planning and execution of corrosion testwork from laboratory scale to pilot scale as well as industrial field trials. Based on this process RHI Magnesita is able to support customers with the most viable refractory lining solution for their processes. The presented recent developments in HF-ITO test set-up show that this corrosion test is a versatile method to evaluate refractory brand qualities under process linked conditions.

References 1. Reuter M et al (2013) Metal recycling opportunities, limits, infrastructure. UNEP International Resource Panel, ISBN 978-92-807-3267-2, 30 2. Gregurek D, Ressler A, Reiter V, Franzkowiak A, Spanring A, Prietl T (2013) Refractory wear mechanism in the nonferrous metal industry: testing and modeling results. JOM 65 (11):1622–1630 3. Malfliet A, Lotfian S, Scheunis L, Petkov V, Pandelaers L. Jones PT, Blanpain B (2014) Degradation mechanism an use of refractory linings in copper production processes: a critical review. J Eur Ceram Soc 34:849–876

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4. Oprea G (2004) Failure mechanism of refractory linings for non-ferrous flash smelting furnaces. In: Proceedings of Tehran international conference on refractories, Tehran, 4–6 May 2004 5. Prestes E, Chinelatto ASA, Resende WS (2009) Post mortem analysis of burned magnesia-chromite brick used in short rotary furnace of secondary lead smelting. Ceramica 55:61–66 6. Wei L (2000) Corrosion of refractories in lead smelting reactors. MSc thesis, University of British Columbia 7. Barthel H (1981) Wear of chrome magnesite bricks in copper smelting furnaces. Interceram 30:250–255 8. Yazawa A (1974) Thermodynamic considerations of copper smelting. Can Metall Q 13 (3):443–453 9. Fontains L, Coussement M, Maes R (1980) Some metallurgical principles in the smelting of complex materials, pp 13–23 10. Bale CW, Chartrand P, Decterov SA, Eriksson G, Hack K, Mahfoud BR, Melançon J, Pelton AD, Petersen S (2002) FactSage thermochemical software and databases. Calphad J 62:189–228 11. Gregurek D, Ressler A, Reiter V, Franzkowiak A, Spanring A, Drew B, Flynn D (2013) State of the art refractory test work for the nonferrous metals industry, Warrendale, TMS

Investigation of Refractory Failure in a Nickel Smelting Furnace Wilson Pascheto, Roy Berryman, Robert Beaulieu and Maysam Moham

Abstract This paper presents results of investigations performed as part of a Root-Cause-Failure-Analysis (RCFA) of a DC smelting furnace that was brick lined with MgO-based refractories. An unusual continuous contraction of the furnace hearth was identified during operation. Cracks in the refractories were envisaged after a major shutdown of the furnace. To determine the mechanism and causes of the unusual contraction behaviour of the furnace hearth, refractory samples were taken from the furnace at different locations and subjected to different analysis including Scanning Electron Microscopy (SEM), Energy-dispersive X-ray spectroscopy (EDX), QEMSCAN and Electron Probe Micro Analysis (EPMA). Creep testing of refractories under different loads (ASTM C832) was also employed to evaluate the mechanical integrity of the used bricks in comparison with the new ones. Results of the various characterisation analyses revealed that there was an interaction between refractory bricks and molten material during furnace operation, which resulted in reduced creep resistance of the MgO-based refractories. This was verified by comparing creep performance of the used brick samples and new brick samples at different temperatures and applied stresses. Keywords Refractory Creep



Furnace integrity



Failure analysis

Introduction Nickel is a versatile metal with a variety of industrial applications including as an alloying element in stainless steels, and in electroplating, storage batteries, shape memory alloys, superalloys, and as an alloying element for non-ferrous metals and W. Pascheto ⋅ M. Moham (✉) XPS Expert Process Solutions, Falconbridge, ON P0M 1S0, Canada e-mail: [email protected] R. Berryman ⋅ R. Beaulieu Koniambo Nickel SAS, BP 679 Voh, NC, France © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_15

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other corrosion resistant alloys [1]. The pyrometallurgical processing of nickel from ores or recycled scrap involve smelting, converting and/or refining furnaces. The reliable operation of these furnaces strongly depends on the refractory materials used in their vessels, which is usually subjected to aggressive process conditions [1]. Resistance of a refractory lining to the combined action of the thermal, chemical and mechanical loads, remains the primary factor to evaluate the refractory [2]. Strange behaviour of a nickel DC smelting furnace in a nickel smelter plant had been reported after installation and during operation of this furnace. Unlike typical brick lined furnaces, the hearth of this furnace exhibited an initial expansion followed by a continuous contraction during its 22 months of operation. This furnace melts the feed (calcined and reduced nickel ore) and produces two liquid streams of slag and crude ferronickel metal. Cross section of the furnace hearth is schematically shown in Fig. 1. The furnace hearth is constructed in an inverted dome-shaped pattern of conductive and non-conductive bricks. An electrical path in the conductive section of the hearth was provided by stainless steel plates in the upper hearth, and magnesia-carbon bricks in the middle and lower hearths. The non-conductive region was provided by magnesia bricks in the upper and lower hearth. A compressive load is provided by a circumferential binding system, comprised of a number of spring-loaded steel segments. This is a unique configuration for such a hearth. After 22 months of operation, the plant experienced a leak of liquid metal from the furnace. This leakage was not caused by the hearth contraction, but the resulted shutdown provided suitable conditions to investigate the root cause of this atypical behaviour, which could have eventually resulted in the furnace failure. This article reports results of investigations and analyses to determine the root cause of the hearth contraction of the DC smelting furnace. In order to conduct this analysis, equipment inspections and sample collections were conducted. Lab analysis techniques including Scanning Electron Microscopy (SEM), Energy-dispersive X-ray spectroscopy (EDX), QEMSCAN, and Electron Probe Micro Analysis (EPMA) were employed to analyze the collected samples.

Fig. 1 Cross section of the furnace hearth

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Inspection and Experimental The refractory bricks at different sections of the cooled furnace hearth were inspected visually and the observations were recorded. During these inspections, some refractory samples were collected for laboratory analysis to determine the failure mechanism, the main cause and also the contributing factors. Besides, operation data was also gathered from the plant control room to complement the analysis results. To evaluate the integrity of the refractory bricks, the collected samples were sectioned, mounted in epoxy and polished to one micron finish. Distribution of elementals in the refractory bricks were analyzed using a Scanning Electron Microscopy (SEM) and Energy-dispersive X-ray spectroscopy (EDX). QEMSCAN and Electron Probe Micro Analysis (EPMA) were employed to determine the composition of different phases in the refractory samples. The primary importance of an EPMA is the ability to acquire precise and quantitative elemental analyses, primarily by wavelength-dispersive spectroscopy (WDS). QEMSCAN, on the other hand, creates surface phase maps through scanning by a high-energy accelerated electron beam. QEMSCAN data includes bulk mineralogy and calculated chemical assays. By mapping the sample surface, textural properties and contextual information such as particle and mineral grain size and shape, material liberation, elemental deportment, porosity, and matrix density can be calculated, visualized, and reported numerically [3, 4]. Some of the collected refractory samples were subjected to creep tests according to ASTM C832 standard (Standard Test Method of Measuring Thermal Expansion and Creep of Refractories Under Load) [5]. In this method, test specimens sectioned from refractory bricks were placed in a furnace and subjected to a prescribed compressive stress. Linear displacement/deformation of the test samples parallel to the direction of the applied stress was recorded continuously as the furnace was heating up. The highest tested temperature was 1550 °C as other material responses such as corrosion, oxidation and sintering could be activated at higher temperatures. Each testing sample was located in a furnace at room temperature and its linear change was measured during heating from room temperature to test temperature. Once the test temperature was reached, the sample was soaked at temperature for 50 h. Only the linear changes during the soaking period are discussed here.

Results Inspection Results The furnace hearth was inspected throughout the demolition to record the condition of the bricks at different sections including upper and lower hearth in the conductive and non-conductive sections. Following is a summary of the furnace hearth inspection results and observations:

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Bricks appearance and dimensions: Although the dimensions of the bricks at the cold side of the vessel were very close to the drawing dimensions, the bricks at the hot face of the furnace hearth were significantly thinner and appeared to be plastically deformed. Moreover, brick measurements revealed that the upper hearth bricks had elongated up to 2% from a nominal length. It was found that the degree of hearth brick deformation was similar throughout the furnace. Discoloration was also observed on many bricks in the conductive and non-conductive sections. It was noted that more discoloration was observed in the samples with higher dimensional changes. During the hearth demolition, disintegration and discoloration of the MgO–C conductive bricks was also observed. This could be an indication of decarburization of the bricks through oxidation. Physical properties: Although the MgO and MgO–C bricks should not have any magnetic properties, some of them exhibited magnetic properties as they were tested by a magnet. This could be an indication of iron impregnation. Hearth Cracking: A horizontal fracture in the upper hearth layer was observed. This crack was seen around the full circumference of the furnace. The crack was fully infiltrated with multiple layers of metal. In some regions, more than one horizontal fracture was observed. Metal infiltration: Metal infiltration was observed in the gaps between bricks at the hot face. The distribution, thickness, and shape of the metal infiltration varied throughout the hearth. The metal infiltration between brick joints was typical and would not affect brick properties if the bricks were not impregnated (diffused). Stainless Steel Plate: Most of the original stainless steel plates between the MgO bricks in the conductive section of the furnace hearth could not be located as they were either entirely removed or indistinguishable from metal infiltration.

Elemental and Phase Analysis (SEM/EDX, QEMSCAN, EPMA) Elemental and phase composition of the refractory brick samples were analyzed using SEM/EDX, QEMSCAN and EPMA on mounted and polished samples. Figure 2 shows an SEM image and EDX results of samples taken from the middle of a MgO brick and at the stainless steel-MgO brick interface. Figure 2a shows the presence of FeNi-rich phase between MgO grains of the brick, indicating infiltration and diffusion of ferronickel molten metal into the brick. It was noted that metal infiltration was more significant on the edges of the MgO bricks. Comparing visual examination results with SEM results revealed that bricks with more metal infiltration exhibited more discoloration and higher deformation. Figure 2b, on the other hand, confirms that the stainless steel sheet at the vicinity of the MgO bricks was replaced by ferronickel and sulfur containing metallic compounds. These results support the visual observation on the absence of stainless steel sheet during inspection.

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Fig. 2 SEM images and EDX results of a MgO brick (a) and the 304 stainless steel-MgO brick interface (b)

Two samples of refractory bricks from conductive and non-conductive sections of the furnace hearth were subjected to QEMSCAN and EPMA analyses. The QEMSCAN and EMPA results of these two samples are presented in Figs. 3 and 4. According to these results, the brick samples contained a significant amount of ferronickel (up to 23 wt%) and FeNi sulphide (up to 7 wt%). This verifies impregnation of ferronickel metal and sulphide into the MgO at different sections of the bricks. These phases are mostly distributed in the grain boundaries, which are the preferred diffusion path for molten metal. It was also noted that diffusion of iron in the MgO grains was more significant in the sample taken from non-conductive section (Fig. 4). It is worth mentioning that metal impregnation was observed on the surface of bricks and also deeper into the bricks. It has been shown that metal impregnation can change refractory properties such as thermal expansion coefficient, thermal conductivity, electrical conductivity, and mechanical strength [6]. Degradation of the stainless steel sheet is also shown in Fig. 3. As can be seen, part of this sheet is broken and lost (top half) and the rest was replaced by a combination of ferronickel, FeNi sulphide and sulphidised steel, consistent with the SEM/EDX results and site observations. This may indicate direct contact between the upper hearth bricks and molten metal.

Creep Tests As aforementioned, the MgO bricks were found deformed at the hot face of the furnace hearth. Change in their dimensions with respect to their drawings and also

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Fig. 3 QEMSCAN and EPMA results of a sample of MgO brick from conductive section

Fig. 4 QEMSCAN and EPMA results of a sample of MgO brick from non-conductive section

in comparison with the bricks installed at the cold side of the brick line could indicate high temperature creep. To determine whether creep performance of the bricks were changed during the furnace operation or not, samples of the impregnated bricks were subjected to creep tests according to ASTM C832 standard [5]. Creep tests were conducted at two different stress levels and two different temperature sequence according to Table 1. The stress levels were selected to simulate the magnitude of stresses applied by the furnace hearth binding system. Creep performance data is limited for most of the commercially available refractories [7, 8]. Therefore, new brick samples were also tested similarly and their results were used as benchmarks to evaluate performance of the used bricks. The orientation of the refractory samples from the furnace hearth was marked during sample collection to identify the sample position in regards to the hearth surface. The position and orientation of test samples are schematically shown in Fig. 5. All creep tests were performed by applying compression with stresses oriented parallel to the furnace hearth, as illustrated in this scheme.

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Table 1 Testing temperatures at different stress levels Compressive stress (Mpa) 0.7 1.5 + tested, − not tested

Testing temperature (°C) 1150 1200 1250 1300 + +

+ +

− −

− +

1350

1400

1450

1500

1550

+ +

− +

− +

− −

+ +

Fig. 5 Position of the creep samples from used refractory bricks and direction of applied compaction stress

Results of the creep tests on used and new brick samples are reported in Fig. 6 for two compressive stress levels of 0.7 and 1.5 MPa. Creep is presented as total irreversible deformation that resulted during the test period. It can be seen that the used and unused bricks showed different creep behaviours. At the highest tested temperature (1550 °C), creep tests of the used samples could not be completed at both stress levels as the samples collapsed prior to reaching the test time of 50 h. Regardless of the test temperature and applied stress, unused brick contracted during creep tests and the contraction increased slightly with temperature within the range of 1150 and 1350 °C, which is in agreement with the creep theory. However, a sharp increase occurred at temperatures higher than 1400 °C. Creep rate of the new bricks at 1550 °C is close to the calculated rate based on the previously reported results for MgO bricks [8]. On the other hand, the used bricks showed expansion when submitted to compressive stress at temperatures up to 1200 and 1350 °C at the applied stresses of 1.5 and 0.7 MPa, respectively. Expansion is expressed by a positive permanent linear change during the creep tests. This is a consequence of the bulk expansion that the used refractories exhibited, which could be attributed to the different physical and mechanical properties of the impregnated metallic phases compared to the MgO refractory grains. Above 1450 °C, the brick contracted quickly due to high creep rate when it was subjected to 1.5 MPa compressive stress. It is worth mentioning that one of the samples tested at 1450 °C didn’t reach the end of the test. None of the used samples tested at 1550 °C could be completed due to excessive brick

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Fig. 6 Total irreversible dimensional change of the used and unused brick samples as a function of temperature, obtained at two compression stress levels of 0.7 MPa (a) and 1.5 MPa (b)

Fig. 7 Creep rate of the new bricks as a function of 1/T (K−1)

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contraction. These results also demonstrated that the creep rates of the used and unused bricks increased with compressive stress, as would be expected. Creep rate and activation energy of the new bricks were calculated by plotting the creep rate as a function of 1/T. The results are shown in Fig. 7. The obtained creep formula is also shown in this figure, where dε/dt, σ, T, and R is creep rate, stress, absolute temperature and gas constant, respectively. Creep rate could not be calculated for the used bricks as they showed expansion at some of the tested temperatures and they failed when tested at 1550 °C.

Discussion In this furnace, expansion of installed bricks and therefore expansion of furnace hearth is constrained by application of compressive stresses using a binding system. It was reported that this furnace showed atypical behaviour as the hearth expanded at the start of operation and then started to contract thereafter. The bricks installed in the furnace hearth are always under compression stress and operate at high temperature, which provides suitable conditions for creep. Site inspection and observations supported hearth brick creep as a hearth contraction mechanism. These evidences include gradual hearth contraction and measured hearth brick deformation. Considering the site observations, stainless steel plate loss and refractory creep can be considered as the hearth contraction mechanisms. However, it is not possible to pick one of these two phenomena as the main mechanism as most of the bricks at the hot face were deformed and also significant number of stainless steel sheets between the bricks in conductive section were damaged. Loss of stainless steel sheets and replacing them by ferronickel, as shown by EDX and EPMA analyses, demonstrated a direct contact between the upper hearth bricks and molten metal. This means that at periods of operation, no frozen heel existed on hearth surface. Lack of frozen heel is attributed to a high molten metal temperature in the furnace. It was shown that the used bricks had lower creep resistance than the unused ones. Lower creep resistance of the used bricks could be attributed to metal impregnation. Based on the laboratory analysis results and the site observations, the MgO bricks in both conductive and non-conductive sections were impregnated by ferronickel and its sulphide products. This diffusion was mostly through the grain boundaries of MgO bricks. Moreover, iron penetrated into the MgO grains, especially in the non-conductive section bricks. Metallic infiltration was also observed in the refractory fillings that were used to fill the gaps between the bricks. The behaviour of refractory materials under high temperature load is highly affected by the presence of low melting point phases. These phases change the grain boundary properties in the refractory brick, especially if it is wetting the grain boundary. At high concentration, the phase with low melting point destruct the direct connections between grains, contributing to a weaker bonds on grain boundaries and lower resistance of the MgO refractory bricks to creep at elevated temperatures [6].

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Site inspection revealed that deformation of the MgO brick near hot face was consistent with the presence of large amount of metal infiltration and the presence of sulphides. This means that creep was more significant when a brick was subjected to a higher level of metal impregnation. Reviewing the historical operating conditions of the furnace reveals that the furnace had been usually operated at molten metal temperature of 1560 °C (tapping temperature). As historical data shows, more than 30% of the metal taps from this furnace were significantly above the design value. It should be noted that the brick temperature at the hot face of furnace hearth could not be measured directly during operation, but only estimated through heat transfer calculations. It is believed that metal penetration is promoted by operating at high temperature as the molten metal diffusion depth is determined by the negative temperature gradient across the refractory lining. In fact, at higher molten metal temperature during operation, molten metal can penetrate deeper into the refractory brick and have more detrimental effect on its performance [6]. Considering the above discussion, high operating temperature of the furnace hearth could be considered as the root cause of the furnace contraction as it contributed to the absence of frozen heel layer, degraded stainless steel sheet and metal impregnation of upper hearth bricks. The latter led to decrease in the creep resistance of the brick and, as a consequence, hearth contraction. Moreover, high operating temperature facilitated the creep deformation as it is a direct function of temperature. Wereszczak et al. [8] introduced maximum 5% creep deformation in 800 test hours as the accepting criterion for bricks in high temperature application under stress. Using this criterion, the MgO bricks used in this furnace should not be used at temperatures above 1500° when they are subjected to 0.7 MPa or higher compressive stresses. It is worth mentioning that brick decarburization, which was detected in the site inspections, could have contributed to higher temperature at upper hearth bricks. Decarburization of bricks in conductive section could lower their thermal conductivities, resulting in higher temperatures of upper hearth refractory bricks.

Conclusions In this article, root cause of the contraction behaviour of a DC smelting furnace was evaluated by collecting the information through site inspections and analyzing the refractory brick samples from different sections of the furnace hearth. SEM/EDX, EPMA and QEMSCAN techniques were employed to analyze microstructure and phase composition of the refractory samples. Refractory samples were also subjected to creep tests at different temperatures and compressive stress levels to evaluate their mechanical properties at high temperature in comparison with the new refractory samples. The main conclusions of this study are as follows:

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• Furnace hearth contraction resulted from the MgO refractory creep under compression and damage in stainless steel layers between the bricks in conductive section of the upper hearth. • Very high operating temperature is believed to be the root cause of the furnace contraction as it activate the following damage mechanisms: – Creep deformation of MgO bricks, by definition, is a function of temperature and higher creep rate is anticipated at higher temperature. – High operating temperature resulted in metal (FeNi) impregnation of MgO bricks, which had detrimental effects on their high temperature creep performance. – High temperature of the molten metal in the furnace prevented formation of a frozen heel above the hearth surface. As a consequence of direct contact with molten metal, many of the stainless steel sheets between the refractory bricks were lost, sulphidised, or replaced by ferronickel. • The creep test results was valuable in establishing a furnace operating strategy by not allowing the metal temperature to reach high values, and therefore cause hearth deformation by creep of MgO refractory. It also provided data to review furnace design, such as binding stresses and furnace cooling.

References 1. Crundwell FK, Moats MSR, Venkoba R, Timothy GD, William G (2011) Extractive metallurgy of nickel, cobalt and platinum-group metals. Elsevier, Amsterdam, Netherlands 2. Malfliet A, Lotfian S, Scheunis L, Petkov V, Pandelaers L, Jones PT, Blanpain B (2014) Degradation mechanisms and use of refractory linings in copper production processes: a critical review. J Eur Ceram Soc 34:849–887 3. Gottlieb P, Wilkie G, Sutherland D, Ho-Tun E, Suthers S, Perera K, Jenkins B, Spencer S, Butcher A, Rayner J (2000) Using quantitative electron microscopy for process mineralogy applications. J Miner Met Mater Soc 52:24–25 4. Jansen W, Slaughter M (1982) Elemental mapping of minerals by electron microprobe. Am Miner 67:521–533 5. ASTM C832 (2015) Standard test method of measuring thermal expansion and creep of refractories under load 6. Schacht C (2004) Refractories handbook. CRC Press-Technology & Engineering, New York 7. Ferber MK, Weresczak AA, Hemrick JG (2006) Compressive creep and thermophysical performance of refractory materials. Final technical report, Office of Energy Efficiency & Renewable Energy, U.S. Department of Energy 8. Wereszczak AA, Kirkland TP, Curtis WF (1999) Creep of CaO/SiO2-containing MgO refractories. J Mater Sci 34:215–227

Mathematical Modeling of Waterless Matte Granulator for Debottlenecking of Conventional Sulfide Smelters A. Navarra and F. Mucciardi

Abstract Within conventional copper and nickel-copper sulfide smelters, ladles of molten matte are fed into converters. These converters blast oxygen-enriched air into the matte, thereby eliminating iron and sulfur. The converting reactions are exothermic and, indeed, the converter heat balance is often a limiting consideration on the smelter throughput. If a portion of the matte were fed in solidified (granulated) form, this would support higher oxygen enrichment, and lower volumes of converter offgas, allowing higher throughput. This approach is not applied in conventional smelters, partly because of the copious amounts of water that must be evaporated in typical granulators, as in the Kennecott-Outotec process. The current paper recalls a waterless matte granulator that had been pioneered in the 1990’s, and is applicable to conventional smelters. A mathematical formulation is presented that can be used in the context of simulation-based optimization, to estimate the size and impact of waterless matte granulation.



Keywords Smelter operations Matte granulation Finite difference Simulation-based optimization





Debottlenecking

Introduction McGill University collaborated with the Noranda Technology Center (NTC) in the 1990’s to develop a matte granulation device that could increase the throughput of conventional copper and nickel-copper smelters [1, 2]. The device incorporated the concepts of heat pipes, inasmuch as a working fluid evaporates from a hot surface, and condenses onto cooling tubes before returning to the hot surface, so that the heat is transported in latent form. The most novel aspect of the device was that it did A. Navarra (✉) Universidad Católica Del Norte, 0610 Angamos, Antofagasta, Chile e-mail: [email protected] F. Mucciardi McGill University, 3610 University Street, Montreal, QC H3A 0C5, Canada © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_16

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not consume any water, and was therefore called a “waterless matte granulator”. Experiments were successfully performed at the NTC (Fig. 1), but interest eventually waned, partly due to the lack of quantitative tools to analyze the system-wide impact that the device could have on smelter operations. The waterless matte granulator (WMG) has been discussed in recent years [3–5], in the context of smelter debottlenecking and discrete event simulation (DES). Indeed, WMG provides an operational degree of freedom to adjust the heat content entering a conventional converting aisle; by adjusting the balance of molten versus granulated matte (Fig. 2), converting can accommodate a higher oxygen enrichment without risk of overheating [4]. As with Kennecott-Outotec flash converting, the higher oxygen enrichment leads to a stronger offgas that is favorable for acid production [6]. Indeed, an assessment of WMG must consider numerous operations, including smelting, converting, oxygen enrichment, offgas handling, as well as other critical aspects that may be indirectly related (availability of different classes of cold charges, flexibility to handle minor elements, etc.). Discrete event simulation is especially well-suited for analyzing the potential impact of WMG, since it supports extensible dynamic models and submodels that can be developed in increasing levels of detail [5]. Even in the early stages of a debottlenecking project, a DES framework can be configured to quantify the overall productivity and sustainability metrics of a smelter, over hundreds of days of operation. As the project advances, increasingly precise data is available to detail the most critical aspects within the DES framework, thereby connecting WMG to these system-wide metrics. Simulation-based optimization is often applied in conjunction with DES [7], as well as finite difference (FD) and finite element (FE) methods, and potentially other computational approaches [8–10]. In particular, an optimization procedure based on Mesh-Adaptive Direct Search can optimize smelter operational policies at the level of a DES framework, while simultaneously adjusting the design parameters of a waterless matte granulation system, at the level of a finite difference scheme. The main contribution of this paper is indeed to develop a computationally efficient

Fig. 1 Experiments performed at the Noranda Technology Center (adapted from [2]), a WMG roll with both end-plates detached, b WMG roll with one end-plate detached, c WMG in operation

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Fig. 2 Debottlenecking of conventional smelter using waterless matte granulation

Fig. 3 Simulation-based optimization for design and evaluation of incremental technological changes, such as installation of WMG in a sulfide smelter

finite difference scheme that is compatible with a DES framework, using the previous work of Mucciardi et al. [1] as a starting point. The new FD method is to approximate the operational rate of heat extraction and matte granulation, as required by the DES framework. Moreover, the FD scheme is to relate the operational data to the geometrical design parameters of the granulator (Fig. 3). Upon completion of the optimization, the geometrical parameters are then subjected to a computationally intensive finite element procedure. The efficient FD scheme is thus a proxy for an intensive FE procedure, within the optimization. In future formulations, the FD scheme may be replaced by a low-resolution FE scheme. Similar approaches have been developed to design aircraft components [9], cardiac prosthetics [10], and other devices. Waterless matte granulation offers a remarkably gradual approach for smelter debottlenecking. Indeed, a pilot WMG can be installed onsite without an urgent need to change the operation of the rest of the smelter; the balance of molten versus granulated matte feed into the converters can then be altered gradually, in conjunction with other parameters. Nonetheless, smelters are reluctant to test any new technology without a means to quantify the potential benefits.

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Customized Finite Difference Scheme For simplicity, the NTC experiments were performed using a single-roll version of the WMG (Fig. 1), although the original concept is to use twin rolls, as illustrated in Fig. 4. Key aspects that permit the successful functioning of the WMG include the wicker interior lining of the rolls that carry liquid sodium (working fluid) to the hot regions, and the flow modifiers within the cooling tubes that ensure effective mixing and warming of the passing air [1, 2]. To quantify the performance of the WMG, Mucciardi et al. [1] developed an ingenious FD scheme to describe the heat extraction through the outer steel shell of the rolls; this scheme couples the rotational speed to the spatial resolution of the sectors (Fig. 5). One time-step Δt is the duration required for a sector to advance into the position of the adjacent sector. For example, if each roll is decomposed into n = 36 sectors, each corresponding to 10°, then each time-step corresponds to a rotation of 10°; thus it requires 36 time-steps for the rolls to complete a 360° rotation and for the sectors to retake their initial positions. Moreover, the formulation considers that each sector position could be in one of two states, depending on whether the outer surface is exposed to ambient air or covered by matte. Without loss of generality, the sector positions are indexed in the direction of the rotation, from 1 to n, such that the first nAir positions are exposed to air, and the remaining positions (n − nAir) positions are covered by matte. Additionally, the indices are circular, such that the successor to sector position i = n is understood as (i + 1) ≡ 1, and likewise the predecessor to position i = 1 is understood as (i−1) ≡ n. Adapting the approach of Mucciardi et al. [1], each sector consists of three nodal temperatures, hence a total of 3n unknowns that are to be solved by the FD scheme,

Fig. 4 Concept of twin-roll waterless matte granulator (Adapted from [3])

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Fig. 5 Illustration of sector position i

namely the inner surface temperature TiInner , the midpoint temperature TiMid , and the outer surface temperature TiOuter . (The original work considered the thickness of the solidified matte crust diCrust to be an additional unknown for the matte-covered sectors, but this has now been eliminated by substitution). The following two equations address both the air-exposed and the matte-covered sectors, representing the heat balance surrounding the inner surface and the middle section of sector i. The lefthand terms describes the transfer of heat through a combination of conduction and convection, whereas the righthand describes the accumulation of heat.  k

    2 2  Inner    π r3 − r2 2A1  Mid Ti − TiInner −1 Ti − TiInner − hInner A0 TiInner − T WF = ρCP r3 − r0 Δt n

ð1Þ  k

      2 2  Mid   π r2 − r1 2A2  Outer 2A1  Mid Ti − TiMid −1 − TiMid − k Ti Ti − TiInner = ρCP r3 − r0 r3 − r0 Δt n

ð2Þ for i = 1 to n, in which k, ρ and CP are the thermal conductivity, density and heat capacity of the steel roll, respectively, Aj is the interfacial area at radial distance rj, TWF is the temperature of the working fluid and hInner is the heat transfer coefficient for the evaporation at the inner surface of the roll. The outer surfaces feature different boundary conditions, depending on whether the sector is air-exposed or matte-covered. Firstly, the air-exposed sectors feature radiation and convection, such that     ησFi A3 TiEnv 4 − TiOuter 4 + hOuter A3 TiEnv − TiOuter i    2A2  Outer −k − TiMid Ti r3 − r0   2 2  Outer  π r1 − r0 Ti − TiOuter −1 = ρCP n Δt

ð3Þ

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for i = 1 to nAir, in which η is the emissivity of the roll, σ is the Stefan-Boltzmann are the view factor, environment temperature and constant, and Fi, TiEnv and hOuter i heat transfer coefficient, respectively, applicable to sector position i. The heat accumulation surrounding the outer surface of a matte-covered sector is given by    2 2  Outer   π r0 − r1 2A1  Outer Ti − TiOuter Mid −1 −k − Ti = ρCP Ti n r0 − r3 Δt 

q̇Cond i

ð4Þ

is the heat conduction rate through the for i = (nAir + 1) to n, in which q̇Cond i solidified matte crust covering sector i. As described by Mucciardi et al. [1], q̇Cond is i quadratically related to the thickness of the crust diCrust ; nonetheless, the current formulation utilizes the linearity of Eq. 4, as diCrust can be solved a posteriori, after q̇Cond has been determined by the FD computations. However, some assumptions i are necessary regarding the profile of the solidifying matte, which are different for the single and twin-roll configurations. From a computational standpoint, the most complicating feature of the finite difference scheme is the fourth power TiOuter 4 that is introduced in Eq. 3, as part of the radiation condition. This complication was not fully described by Mucciardi et al. [1], but is the main focus of the rest of this section. The proposed scheme is to utilize the linearity of Eqs. 1, 2 and 4, while addressing the nonlinearity of Eq. 3. Firstly, Eq. 3 can be re-expressed as      ρCP π r12 − r02 2A2 Outer Δt ⋅ TiMid Ti − 1 + k n r3 − r0       ρCP π r12 − r02 2A2 Outer − hi A 3 + k + Δt ⋅ TiOuter ð5Þ n r3 − r0   4 Env = − ησFi A3 TiEnv + hOuter A T Δt + ησFi A3 Δt ⋅ ui 3 i i placing the linear terms on the lefthand side, and applying the substitution ui = TiOuter 4 , or equivalently TiOuter = u1i

̸4

ð6Þ

for i = 1 to nAir. The linear coefficients of Eqs. 1, 2, 4 and 5 form an invertible matrix, so that the unknown temperatures can be expressed as a linear combination of u = ½u1 u2 ⋯unAir T . In particular, nAir

TiOuter = ci0 + ∑ cij uj

ð7Þ

j=1

in which the coefficients cio and cij can be determined by applying matrix inversion (i.e. Gaussian Elimination) to Eqs. 1, 2, 4 and 5; these coefficients carry the result

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of the linear analysis, essentially, so that the remainder of the formulation can address the nonlinearity. Indeed, Eq. 6 is satisfied if and only if   nAir 1 ̸4 ci0 + ∑j = 1 cij uj = ui , which forms a system of nAir nonlinear equations, nAir

f ðuÞ = ci0 + ∑ cij uj − u1i ̸4 = 0

ð8Þ

j=1

for i = 1 to nAir, that can be solved by Newton’s Method, as described below. After u has been determined, Eq. 6 is used to compute TiOuter for i = 1 to nAir, and Eq. 7 is used to for i = (nAir + 1) to n. Given TiOuter , Eqs. 3 and 4 are then used to compute TiMid , and finally Eq. 2 is used to compute TiInner . Equation 8 describes a system of nAir nonlinear equations, which can be used to solve nAir unknowns that constitute the vector u. To solve for u, the kth iteration of Newton’s Method is given by   3 f1 uk − 1      − 1 6 f 2 uk − 1 7 6 7 uk = uk − 1 − J u k − 1 4 5 ⋮k − 1  fnAir u 2

ð9Þ

in which J ðuÞ denotes the Jacobian matrix [11], whose entries ij are given by ∂fi u − 3 ̸4 δij ðuÞ = cij − i ∂uj 4 in which δij denotes the Kronecker delta. The Jacobian matrix can thus be expressed as 

u−3 J ðuÞ = C − Diag i 4  in which Diag

ui− 3 4

̸4

̸4 

 is the diagonal matrix whose entries are given by

ui− 3 4

̸4

.

Equivalently, the Jacobian matrix can also be expressed as nAir

J ð uÞ = C − ∑

i=1



ui− 3 4

̸4 

ei eTi

ð10Þ

in which ei is the ith column of the nAir × nAir identity matrix. The performance of Newton’s Method depends on the efficient computation of the inverse Jacobian (Eq. 9). Indeed, Eq. 10 presents a particular structure that motivates the sequence

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 J l ð u1 , u2 , . . . , ul Þ = J l − 1 ð u1 , u2 , . . . , ul − 1 Þ −

ul− 3 4

̸4 

el eTl

ð11Þ

such that J0 = C and JnAir ðuÞ = J ðuÞ, which is subject to the Sherman-Morrison Identity [12]. With the understanding that Jl is a function of (u1,u2,…, ul), the identity can be adapted to give ð Jl Þ

−1

= ð Jl − 1 Þ

−1

! ul− 3 ̸4 1 + ðJl − 1 Þ − 1 el eTl ðJl − 1 Þ − 1 4 1 − eTl ðJl − 1 Þ − 1 el

ð12Þ

in which eTl ðJl − 1 Þ − 1 el refers to the lth diagonal entry of ðJl − 1 Þ − 1 ; also, ðJl − 1 Þ − 1 el eTl ðJl − 1 Þ − 1 is a matrix whose ij entry is the product of the il and lj entries of ðJl − 1 Þ − 1 . Therefore, supposing that ðJ0 Þ − 1 = ðC Þ − 1 has been computed and stored, the repeated application of the Sherman-Morrison Identity incorporates the effect of u1, the effect of u2, and so on, until all of u has been incorporated, finally obtaining ðJnAir ðuÞÞ − 1 = ðJ ðuÞÞ − 1 , which is applied in Eq. 9. There are thus nAir Sherman-Morrison iterations (Eq. 12) within each Newton iteration (Eq. 9). In summary, the linear coefficients of Eqs. 1, 2, 4 and 5 are used to obtain the coefficients cio and cij of Eq. 7, the latter of which is used to construct an nAir nAir matrix C whose inverse is computed and stored; the inverted matrix C−1 is then employed within a sequence of Newton iterations (Eq. 9) with nested Sherman-Morrison iterations (Eq. 12), to give values for TiOuter for i = 1 to nAir, by way of Eq. 6. The remaining temperatures are given by the successive application of Eq. 7, Eqs. 3 and 4, and finally Eq. 2. Following a similar approach taken by [5], this FD scheme can be implemented firstly in Microsoft Excel and/or Matlab, before being implemented in Rockwell Arena© as part of a DES framework. One aspect which admittedly is not well described in the current paper is the reprein Eq. 4, which will be described in a future paper. Typical values sentation of q̇Cond i for heat transfer coefficients and working substance temperatures were obtained by Mucciardi et al. [1, 2], and validated by the experimental trials at the Noranda Technology Center.

Conclusions and Future Work WMG is a particular technology that can have major impacts in conventional copper and nickel-copper smelters; the development of computational tools to evaluate these potential impacts may be an important step toward reviving the earlier interests that had been strong in the 1990s. A following paper will present sample computations that incorporate the FD scheme within a DES framework. Meanwhile, there are ongoing efforts to build relationships with industrial partners, so that the experiments performed at the Noranda Technology Center can be

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repeated, and extended to the pilot- and production-scale. The ongoing issue of effective SO2 capture is related to the converter heat balance, which is likely to be a point of interest for WMG. Moreover, the computational approaches developed for WMG could be replicated for other technologies. Beyond the immediate economic drivers for smelters to meet increasingly stringent environmental regulations, and for antiquated smelters to be competitive with newer smelters, there is an intergenerational requirement to modernize the education of extractive pyrometallurgy. If the WMG experiments can be repeated within industrial laboratories, they may eventually be repeated in universities, as part of undergraduate or graduate laboratory exercises. WMG is indeed central to the operation of copper and nickel-copper smelters, since it addresses the interface between smelting and converting (Fig. 2). An understanding of how WMG influences the heat balance for the converters and, in turn, how the heat balance is related to oxygen enrichment and offgas handling, can foment a system-wide perspective that is an essential part of modern engineering education [13]. Thus, in parallel to the computational developments and the establishment of industrial partners, the authors also seek education-minded partners who are interested in preparing the next generation of pyrometallurgists.

References 1. Mucciardi F, Jin N, Palumbo E (1998) Modelling a novel, air-cooled, twin roll caster. In: Paper presented at the 37th Annual conference of metallurgists of CIM, Calgary, Alberta, 16– 19 Aug 1998 2. Mucciardi F, Palumbo E, Jin N (1999) A waterless caster for matte/slag granulation. In: Paper presented at the 4th International copper conference, Phoenix, Arizona 10–13 Oct 1999 3. Navarra A, Mucciardi F (2015) Discrete event simulation to quantify upgrades of Peirce-Smith converting aisles. In: Paper presented at the 37th International symposium of application of computers and operations research in the mineral industry, Fairbanks, Alaska, 23–27 May 2015 4. Navarra A, Kuan SH, Parra R, Davis B, Mucciardi F (2016) Debottlenecking of conventional copper smelters. In: Paper presented at the 6th International conference on industrial engineering and operations management, Kuala Lumpur, Malaysia, 8–10 Mar 2016 5. Navarra A, Marambio H, Oyarzún F, Parra R, Mucciardi R (2016) System dynamics and discrete event simulation of copper smelters. Miner Metall Process 34(2):96–106 6. Kojo I, Lahtinen M, Miettinen E (2009) Flash converting – Sustainable technology now and in the future. Paper presented at the 139th TMS annual meeting, San Francisco, California, 15–19 Feb 2009 7. Kleijnen J, Wan J (2007) Optimization of simulated systems: optquest and alternatives. Simul Model Pract Theory 15(3):354–362 8. Audet C, Denis J Jr, Le Digabel S (2012) Trade-off studies in blackbox optimization. Optim Meth Softw 27(4–5):613–624 9. Tribes C, Dubé J-F, Trépanier J-Y (2005) Decomposition of multidisciplinary optimization problems: formulations and applications to a simplified wing design. Eng Optim 37(8):775– 796 10. Marsden A (2014) Optimization in cardiovascular modeling. Annu Rev Fluid Mech 46:519– 546

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11. Burden R, Faires J, Burden A (2016) Numerical solutions of nonlinear systems of equations, Chapter 10. In: Numerical analysis. Cengage Learning, Boston, Massachusetts, pp 641–683 12. Meyer C (2000) Inverses of sums and sensitivity, Section 3.8. In: Matrix analysis and applied linear algebra. Society for Industrial and Applied Mathematics, Philadelphia, Pennsylvania, pp 124–131 13. Crawley E, Malmqvist J, Östlund S, Brodeur D, Edström K (2014) Adapting and implementing a CDIO approach, Chapter 8. In: Rethinking engineering education—the CDIO approach. Springer, New York, pp 181–207

Desulfurization of the Non-ferrous Smelter Flue Gases Based on Scrubbing with a Carbonate Eutectic Melt and Natural Gas Regeneration Valery Kaplan, Nurlan Dosmukhamedov and Igor Lubomirsky

Abstract Sulfur emission in the form of SO2 in flue gases is one of the most serious atmospheric pollutants associated with coal combustion and non-ferrous metals production. The carbonate eutectic method for removing SO2 from flue gases at 723–923 K was initially proposed in the 1970s but despite its great efficiency (SO2 concentration in the flue gas after purification reached 0.003 volume %), it could not be implemented by industry due to the complexity of the carbonate melt regeneration stage. Earlier we proposed a method suited to coal-firing power stations where the melt was regenerated using CO as a reducing agent. However, most metallurgical plants do not use coal and therefore lack a large source of CO. Here we propose a method for removing sulfur from the carbonate eutectic melt by purging it with natural gas or a natural gas/air mixture, which are available in the vast majority of metallurgical plants. This reaction leads to the reduction of sulfate to H2S gas that leaves the melt. The experiments we conducted show that nearly complete sulfur removal from the melt is possible at 823 K and that the reaction rate is sufficiently high for a large scale process. One can foresee that this carbonate melt-based SO2 removal technique may become a practical and economically attractive method for limiting sulfur emission to the atmosphere from non-ferrous metallurgical processing plants.





Keywords Sulfur emission Environment Natural gas Carbonate eutectic melt Flue gases Non-ferrous metallurgy





V. Kaplan (✉) ⋅ I. Lubomirsky Weizmann Institute of Science, Rehovot, Israel e-mail: [email protected] N. Dosmukhamedov Kazakh National Research Technical University After K.I. Satpayev, Almaty, Kazakhstan © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_17

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Introduction Mines and metallurgical plants produce large amounts of waste because the ore constitutes only a small fraction of the total volume of the mined material. In the metallurgical industry, production of Cu, Pb, Ni, and Zn causes the greatest damage to the environment [1–3]. Gases emitted with modern smelting technology (10–30 volume % SO2) can be processed to elemental sulfur at a ratio of one ton of sulfur for each ton of metal [4]. Production of one ton of non-ferrous metal by traditional, older processes produces flue gas with a relatively low SO2 content ( 760 ◦ CÞ SO3 + MgO → MgSO4 ð < 1050 ◦ CÞ SO3 + CaO → CaSO4

As described in the second post-mortem study As-oxide can react as As-pentoxide As2O5 with volatilized lead (PbO) in the off-gas. These phases become liquid >640 °C depending on the As2O5/PbO-ratio (Fig. 5).

Fig. 5 PbO-As2O5 phase diagram [15]

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Fig. 6 Formation of As-containing solid phases (at a temperature of 1200 °C and pO2 = 1*10−9) in different MgCr-refractories (bricks A–F), i.e. newly formed reaction product of As-containing process phases with brick phases

With help of the scanning electron microscope the PbO-As2O5 phases Pb8As2O13 and Pb4As2O9 have been detected. Following reactions can be assumed: 8PbO + As2 O5 → Pb8 As2 O13 4PbO + As2 O5 → Pb4 As2 O9 In addition to the mineralogical investigation FactSageTM calculations with data bases FToxide and FTmisc were carried out to study the theoretical corrosion of the brick components as well as phase formations. The calculation results were compared with the microscopic investigation results. As-phases react with the MgO component, forming Mg3(AsO4)2 and hence represent an additional corrosion factor for the refractory bricks. Figure 6 shows several magnesia-chromite bricks with different chemical composition reacting with an As-containing slag. Brick F forms the lowest amount of solid phase. Due to database limitations, an additional reaction with sulfur and formation of a mixed Mg-As-sulfate as mineralogically detected in the post mortem sample cannot be seen in the calculations.

Conclusion As the arsenic content in copper ores has become an increasingly important factor over the last years it is essential to know the influences of arsenic in the copper smelting process and its contribution to refractory wear. Especially post-mortem studies are important for industrial application and specific customer requirements.

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The investigated magnesia-chromite brick out of a copper smelting furnace shows mainly continuous wear by hot erosion and is severely attacked by Fe-silicate slag (fayalite type Fe2SiO4). Enrichment of arsenic caused severe corrosion of mainly the magnesia component forming Mg-As-sulphate as main reaction product. Arsenic attack on the refractory material is not the main chemical attack, but represents an important additional factor influencing lining wear and lifetime. In the second post-mortem study on the magnesia-chromite castable arsenic reacted with lead under formation of Pb-arsenate of type Pb8As2O13 and Pb4As2O9. Two corrosion mechanisms were observed: Chemical attack by SiO2 on the magnesia under formation of forsterite. Chromite was also highly attacked by Sb-Feand Sn-oxide. Additionally the castable got deeply infiltrated and corroded by Pb-arsenate. The observed attack on the magnesia-chromite brick and castable by As compounds in the zone of liquid metal and slag (for the brick) and off-gas area (for the castable) may have been caused by insufficient As volatilization due to not optimized process conditions. FactSageTM calculations, based on available data, show that As will mainly attack the basic component of the refractory under formation of e.g. magnesium-arsenate Mg3(AsO4)2. Arsenic can also react with other components in the off-gas e.g. lead, forming lead-arsenic oxide phases, which can infiltrate and corrode the microstructure. Post-mortem investigations on used refractory materials represent an important prerequisite for the product development, as well as for special engineered lining concepts to support our customers.

References 1. Larouche P (2001) Minor elements in copper smelting and electrorefining 2. Routschka G (2011) Refractory materials, 4th edn. Vulkan-Verlag, Essen, p 505 3. http://www.conductivity-app.org/single-article/cu-overview, European Copper Institute, Accessed 26 Feb 2018 4. Hoffmann JE (1993) Remediating copper smelter dusts: the arsenic problem. JOM:30–31 5. Piret NL (1999) The removal and safe disposal of arsenic in copper processing. JOM:16–17 6. Imris I, Klenovcanova A (2003) Removal of arsenic, antimony and mercury from copper concentrates by roasting process. In: Proceedings of copper 2003, Santiago, Vol IV, pp 125–139 7. Dalewki F (1999) Removing arsenic from copper smelter gases. JOM:24–26 8. Lindkvist G, Holmström A (1983) Roasting of complex concentrates with high arsenic content. Adv Sulfide Smelt:451–472 9. Alvear GR et al (2008) Copper isasmelt—dealing with impurities. In: Proceedings of sohn international symposium advanced processing of metals and materials, Vol. 8, International symposium of sulfide smelting, TMS, 2006, pp 673–685 10. Wilkomirsky I, Parra R, Parada F, Balladares E (2013) Physico-chemistry and kinetics mechanisms of partial roasting of high-arsenic copper concentrates. In: Proceeding of copper 2013, Chile, pp 539–552

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11. Weisenberg IJ, Bakshi PS (1979) Arsenic distribution and control in copper smelters. J Metals:38–44 12. Grund SC, Hanusch K, Wolf HU (2008) Ullmann’s encyclopedia of industrial chemistry. Arsen Arsen Compd 13. Holleman AF, Wiberg N (2007) Lehrbuch der anorganischen chemie, Auflage. Walter de Gruyter & Co, Berlin, p 102 14. Gregurek D, Reinharter K, Majcenovic C, Wenzl C, Spanring A (2015) Overview of wear phenomena in the lead processing furnaces. J Eur Ceram Soc 35(6):1683–1698 15. Firoozi S (2005) Thermodynamics and mechanisms of lead softening: 16–24

Improved Copper Smelter and Converter Productivity Through the Use of a Novel High-Grade Feed Eugene Jak, Denis Shishin, Will Hawker, James Vaughan and Peter C. Hayes

Abstract Copper sulphide processing technologies face increasing pressures associated with decreasing concentrate grade leading to increasing thermal inefficiency and lower productivity. Impurity concentrations are on average increasing, creating potential environmental risk and additional treatment costs. In copper flash smelters dust, partially oxidised materials and fume formed from the condensation of volatile impurities, are routinely recycled to the feed. In the converting stage the heat balance is maintained by charging anode reverts and other inert materials. In both cases, the thermal energy available from sulphide oxidation is not fully utilised or optimised. The productivities of both smelter and converter stages can be potentially increased through the addition of a high copper, low iron, low impurity precipitated copper product. Calculations are carried out for fayalite smelter and calcium ferrite converter slags using an optimised FactSage thermodynamic database. The potential for significant increases in smelter and converter productivities using existing technologies are predicted. Keywords Copper smelter



Copper converter



Productivity

Introduction The average grades of copper sulphide ores continue to decline as the high-grade deposits currently utilised for primary copper metal production are depleted [1]. This poses increasing concern on a number of fronts. As ore grades decrease the quantity of ore that must be mined and processed increases inversely with the ore grade; all of this ore produces commensurate quantities of low-grade, sulphide wastes that must be disposed of appropriately so that there is no environmental impact in the short or long term. E. Jak ⋅ D. Shishin ⋅ W. Hawker ⋅ J. Vaughan ⋅ P. C. Hayes (✉) Metallurgical Engineering, School of Chemical Engineering, The University of Queensland, Brisbane Q4072, Australia e-mail: [email protected] © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_21

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The separation of sulphide minerals from the ores requires extensive expenditure of energy in the size reduction or comminution stages of concentrate feed preparation. Studies have shown that even in current operations the greatest energy required in copper metal production via the conventional sulphide flotation/smelting/converting/pyro- and electro-refining route is in the comminution stage [2]. Comminution is usually undertaken at remote sites where there are no connections to major electricity grids and low cost power. Therefore, power must be generated on site, requiring additional capital cost in the provision of infrastructure. These stand-alone operations also use hydrocarbon fuels thus contributing to greenhouse gas emissions. Decreasing ore grade will lead to increased energy consumption and cost associated with the preparation of concentrates from these ores. Low concentrate grades to the smelter also increase the slag volumes created in the smelting stage as the iron associated with the copper concentrate reports to the slag phase, and the slag compositions must be adjusted by flux additions to obtain appropriate physical and chemical properties. Increased slag volumes lead to increased copper losses through dissolved and entrained copper [2]. Hydrometallurgical processes have the advantage that fine grinding of the ore is generally not required in order to expose the copper-containing minerals and enable selective dissolution through variety of leaching technologies. Dissolution of copper from copper oxide based ores is generally readily achieved using dilute sulphuric acid solutions. Selective leaching of copper from sulphide minerals is more difficult but technologies are now being developed to enable this to be undertaken on low grade sulphide ores at industrial scale [3]. The aim of the synergistic copper process concept is to take advantage of the inherent benefits from both the hydrometallurgical and pyrometallurgical process routes [4]. The simplified process flow diagram of the synergistic copper process is shown in Fig. 1. In the proposed process the selective extraction of copper via the low-energy leach of copper oxide or sulphide ores is undertaken, the solution is treated to further separate copper from iron and other impurities by precipitation and then precipitate a solid intermediate copper product. The precipitate can be fed directly into the smelter or pre-calcined using waste heat gas and introduced into the convertor. The application of the process concept is not limited to a particular technology; so in principal it can be used in both flash and bath smelting and converting operations. The proportion of high-copper, low-iron precipitate introduced is determined by the excess enthalpy available from the exothermic reactions taking place in each of the pyrometallurgical smelting and converting reactors. Preliminary calculations have been undertaken using simplified thermodynamic models to estimate the extent to which the precipitate can be added to the smelter feed through blending with the concentrate [5, 6]. These initial calculations have been extended in the present paper to take into account associated converter operations and the recycling of converter slag.

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Fig. 1 Schematic process flow diagram of the synergistic copper process [4]

Methodology The calculations provided in the present are focussed on the use of the top submerged lance technology for both smelting and converting processes. The particular examples selected are the IsaSmelt and Isa Convert processes for copper. The details of the calculation procedure are provided elsewhere [7]. Briefly, this involves calculating both heat and mass balances in both smelter and converter reactors. The enthalpy available in the smelting and converting processes for a given base case, assuming a fixed concentrate feed rate and composition, and calculating the total enthalpy generated in each reactor. The enthalpy loss from each reactor is estimated from plant practice. The differences between these values indicate the excess enthalpy that is available to enable other process adjustments to be undertaken. In the proposed application, this excess enthalpy is used to add high copper, low iron, low impurity precipitated copper product to both smelter and converter. The calculations have been undertaken using FactSage chemical thermodynamic database, which contains thermodynamic descriptions of solid, liquid and gas phases in the pyrometallurgical copper smelting and refining systems, in particular the databases contains solution models for gas species, liquid slags, liquid mattes, liquid metals and solids that are present in these high temperature systems [8]. A process flow diagram illustrating the process is shown in Fig. 2. In the base case a typical copper concentrate containing 25% Cu is selected to obtain a matte grade in the smelter of 70% Cu using a 60% O2 blow; the matte from this step is cooled to room temperature and used as feed to the converter. The smelter slag is discarded and the gas taken to waste heat boiler for heat recovery and separate power generation as is standard practice. The waste heat recovery is not included in

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Fig. 2 Process flow diagram illustrating the process flows incorporated in the copper smelter and converter calculations

the calculations. The cold matte 70% Cu is converted to blister copper assuming 20% Cu reports to the slag and again using 60% O2 enriched gas blow. The resulting converter slag is cooled to ambient temperature and recycled to the smelter to recover copper from the slag. Following the base case calculations using the excess enthalpy, the specific calculations are repeated to include the addition of the precipitated copper product.

Results and Discussion Summaries of the detailed calculations for the base case and with addition of precipitate are given in Table 1, and are discussed in the following paragraphs. The input and output parameters are identified in column #2; individual input and out items are given in the lines and associated units are given in column #3. Base Case The base case calculations (see columns 3 and 4) using 200t h−1 of 25% Cu concentrate and 70% Cu matte show that 8.3 MW is generated in the smelter and 8.7 MW in the converter. Information available from industrial practice indicates that, for similar process conditions, it is estimated that there would be heat losses of approximately 3 MW from the smelting reactor and 6 MW from the converter. These differences primarily reflect the differences in detailed designs since the smelting furnace, having a fayalite–based slag, has a refractory hot face, whereas the converter slag requires the use of freeze lining to contain the aggressive calcium ferrite based slag. The calculations then show for the base case that net energy available for precipitated copper product addition is 5.3 MW in the smelter and 2.7 MW in the converter. Case 1 Maintaining the copper concentrate feed rate at 200t h−1 of 25% Cu concentrate and 70% Cu matte but adding precipitated copper product to the furnaces to obtain enthalpy balances in both smelter and converter, it can be seen from the calculations (see column 5 and 6) that an increase in productivity is obtained from both smelter and converter. The productivity of the smelter is increased from 1260 to 1318 t d−1 and the converter from 1156 to 1273 t d−1; increases of 4.6 and 10.1% respectively. Although approximately equal quantities of precipitated copper product are used in each reactor the benefits are cumulative, resulting in enhanced overall productivity. There are other advantages from the precipitated copper product addition, in addition to the 10% increase in productivity, there is

Conc. Matte Precip. Cu SiO2 flux CaCO3 flux Convert Slag air Tonnage O2

2 Parameter

7 8 OUTPUT 9 Slag 10 Matte 11 Metal 12 Off gas 13 Off gas

INPUT 1 2 3 4 5 6

Col. 1 Line #

d−1 d−1 d−1 d−1 d−1 d−1

t d−1 t d−1 t d−1 t d−1 Nm3 h−1

Nm3 h−1 Nm3 h−1

t t t t t t

3 Conditions/ units

3,345 82,491

3,118 1,762

31,397 34,985

303

964

4,800

4 Base smelter 25% Cu conc.

1,169 1,079 21,233

302

10,262 11,435

109

1,747

5 Base case converter

3,308 81,520

3,128 1,845

30,480 33,964

317

4,800 – 107 965

6 Case 1 Smelter 25% Cu conc. + ppt

1,287 1,142 22,128

318

10,386 11,537

97

– 1,845 101

7 Case 1 Converter + ppt

Table 1 Summary of calculated outcomes of smelter and converter processes for selected scenarios

3,479 84,806

3,564 1,667

31,818 35,455

286

4,800 – 348 1,141

8 Case 2 Smelter 20% Cu conc. + ppt.

1,154 1,031 20,038

287

9,462 10,544

90

– 1,667 73

9 Case 2 Converter + ppt

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(a) No additional fuel required in smelter or converter. (b) No increase in smelter off gas volume. This extra throughput is achieved with the smelter and converter gas volumes of 3,308 t d−1 (81,520 Nm3 h−1 at 35.8% SO2) and 1,142 t d−1 (22,230 Nm3 h−1 at 54.5% SO2) respectively, a total of 4,450 t d−1 (103,750 Nm3 h−1) compared to 3,345 t d−1 (82,490 Nm3 h−1 at 35.6% SO2) for the smelter and 1,080 t d−1 (21,233 Nm3 h−1 at 52.3% SO2) for the converter, a total of 4,425 t d−1 (103,750 Nm3 h−1) for the base case. (c) Reduced tonnage oxygen requirement, approximately 2 vol.% O2 less than base case. From the base case smelter 34,985 Nm3 h−1 and converter 11,435 Nm3 h−1, a total of 46,420 Nm3 h−1, to smelter 33,964 Nm3 h−1 and converter 11,573 Nm3 h−1, a total of 45,537 Nm3 h−1 with precipitate addition. This saving comes from the oxygen associated with the copper, calcium and sulphate in the precipitate. (d) No increase in silica smelter flux requirement. The slag compositions remain essentially unchanged in both smelter and converter practices; within normal operating ranges. (e) Only marginal increases in the smelter slag mass from 3118 to 3128 t d−1, i.e. a 0.3% increase in mass. Case 2 In Case 2, the scenario is examined when there is a decrease in concentrate grade to the smelter; this reflects the trend in the industry of reducing ore and concentrate grade over time. In this example, the copper concentrate feed rate at 200t h−1 of 20% Cu concentrate and 70% Cu matte but adding precipitated copper product to the furnace to obtain enthalpy balances in both smelter and converter. This enables an extra outputs of 178 t d−1 Cu in the smelter and 37 t d−1 Cu in the converter; a total of additional 215 t d−1 Cu on top of a feed of 960 t d−1 Cu in the concentrate. The calculations clearly demonstrate that as concentrate grades decrease the smelter productivities will drop and the available enthalpy will increase. Although this scenario is undesirable, the calculations show the opportunities for compensating for this effect by the addition of precipitated copper product to the smelting and converting reactors. The greater the excess enthalpy the greater the opportunity to add the high copper, low iron precipitate, without major changes to the existing process flowsheet or changes to operating practice. Small additions of precipitated product can be used as feed to smelter, enabling the treatment low grade copper concentrates without the need for pre-preparation, i.e. uncalcined precipitate. In this case minimum capital expenditure is required. However, this scenario does not take full advantage of potential savings and increased throughput available from the use of calcined material that have been illustrated in the examples above. Some capital expense would be required to partially calcine the precipitated copper product to remove chemically and physically bound water; however, this process could be carried using the low grade waste heat gases produced by smelting and converting processes. This is low temperature heat and should complement the existing heat recovery systems in the smelter and improve overall energy utilisation of the plant. The enthalpy requirements for

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different extent of calcination have been previously reported by the authors [6]. Optimum operating practice will differ between operations depending on the characteristics of each smelting operation.

Summary Thermodynamic calculations have been undertaken to predict the outcomes from copper smelting and converting operations using the IsaSmelt and IsaConvert top submerged lance processes. The case studies have demonstrated that excess thermal energy is available in both smelting and converting stages. The excess energy can be usefully used to increase smelter and converter throughput without major capital expenditures or increased operating costs through the addition of a high copper, low iron, low impurity precipitated copper product. Significant increases in productivity are predicted; these opportunities for supplementary feed increase as concentrate grade to the smelter decreases. This offers opportunities to concentrate suppliers and smelters to increase resource utilisation, reduce energy costs and increase existing smelter and converter productivities.

References 1. Mudd GM, Weng Z, Jowitt SM (2013) A detailed assessment of global Cu resource trends and endowments. Econ Geol 108:1163–1183 2. Schlesinger ME, King MJ, Sole KC, Davenport WG (2011) Extractive metallurgy of copper, 5th edn. Elsevier 3. Ekenes JM, Caro CA (2013) Improving leaching recovery of copper from low-grade chalcopyrite ores. Miner Metall Process 30(3):180–185 4. Hawker W, Vaughan J, Jak E, Hayes PC (2017) The synergistic copper process concept, Mineral Processing and Extractive Metallurgy Trans IMM C, 2017, C1-11 5. Hawker W, Vaughan J, Jak E, Hayes PC (2016) The synergistic copper process—a new process route for low-energy copper production, Copper 2016, Kobe, HY-5-4 6. Hawker W (2015) PhD thesis, The University of Queensland, Brisbane, Australia 7. Jak E, Nicol S, Hidayat T, Shishin D, Hayes PC (2016) The potential for energy savings, increased productivity and recovery in copper smelting, converting and recycling through implementation of experimental and thermodynamic modelling research, Copper 2016, Kobe, paper PY3-5 8. Shishin D, Hayes PC, Jak E (2018) Multicomponent thermodynamic databases for complex non-ferrous pyrometallurgical processes. Sulphide Smelting, TMS-CIM, Ottawa

Semi-discrete Dynamics and Simulation of Peirce-Smith Converting A. Navarra, G. Lemoine, N. Zaroubi and T. Marin

Abstract Peirce-Smith converting (PSC) is applied for roughly 50% of primary nickel and 70% of primary copper production. PSC cycles produce batches of ironfree sulfide matte (or blister copper, in the case of copper smelters), that are subject to further processing. However, the number of cycles that can be performed simultaneously is limited by the offgas handling system. Moreover, PSC suffers from variation in the yield and duration of the cycles. This variation is managed by conventional smelter designs, in which the upstream smelting capacity exceeds the nominal converting capacity; PSC is thus a major bottleneck in conventional nickel and copper smelters. Stabilization and standardization of PSC operations can therefore increase smelter throughput. The current paper presents a discrete event simulation (DES) framework to assist in smelter debottlenecking. It features random number generation to represent cycle variability, and time-adaptive finite differences to represent thermochemical complexity. Preliminary computations are presented.



Keywords Peirce-Smith converting Discrete event simulation Mass balance Heat balance Gibbs free energy balance





Introduction Peirce-Smith converting (PSC) is applied for roughly 50% of primary nickel and 70% of primary copper production, which corresponds to 23 and 1270 kt, priced at 9200 USD/t and 5600 USD/t, respectively [1–3]. Moreover, PSC is often the main A. Navarra (✉) Universidad Católica del Norte, 0610 Angamos, Antofagasta, Chile e-mail: [email protected] G. Lemoine ⋅ N. Zaroubi McGill University, 3610 University Street, Montreal, QC H3A 0C5, Canada T. Marin M4 Dynamics, 1 Younge Street, Suite 1801, Toronto, ON M5E 1W7, Canada © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_22

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bottleneck in conventional copper and nickel-copper smelters, as it sets the tempo for all subsequent operations. Thus improvements in PSC impact the overall productivity and sustainability of smelters. The potential for high impact developments is especially notable, considering the world-wide similarity of conventional smelters; improvements implemented in a specific smelter have been generalized and adapted to numerous other smelters. Such improvements have included the incorporation of tonnage oxygen for blast enrichment, and Gaspé punchers to periodically clear the tuyères [1]. Nonetheless, other potential improvements have merit, such as shrouded oxygen injection and waterless matte granulation [4–6], and have not been widely employed in conventional smelters, perhaps due to the lack of adequate quantitative tools to assess their system-wide benefits. A PSC converting aisle consists of a parallel arrangement of converters, which functions together with the preceding smelting furnace (Fig. 1) to eliminate iron and sulfur from the iron-nickel-cobalt-copper sulfide feed. For nickel-copper smelters, the converted output is iron-free matte that is also known as Bessemer matte [7], and for copper smelters its so-called blister copper [8]; in both cases, the output requires further processing to eliminate deleterious minor elements, to retrieve valuable minor elements, and to cast nickel, cobalt and/or copper products. The oxidized sulfur leaves the bath in the form of SO2, as part of the offgas. This offgas is essentially a mixture of N2 and SO2, but it typically carries some unreacted O2 and considerable quantities of flue dust; for typical studies, it is usually a reasonable approximation to assume a single stable oxidized form of sulfur, namely SO2. In contrast, there are two stable iron oxides, FeO and Fe3O4, which coexist within a fayalite or olivine slag [9]. The balance of FeO and Fe3O4 is quantified by the degree of iron oxidation α = Fe3+/Fe2+ in the slag [10], which is computed through a Gibbs Free Energy balance, as discussed in the following section. Most importantly, α determines how much oxygen is required for each ton of feed, having a given composition; this ultimately determines the production rate of the entire smelter. The current paper focuses on the development of a unified computational framework for nickel-copper and copper smelter analysis, and is part of a third installment (Fig. 2). The first installment was described in the first symposium on the Pyrometallurgy of Nickel and Cobalt in 2009 [5]. This work extended an operational cost model that had originally been developed for copper smelters [11];

Fig. 1 Smelting furnace and converting aisle eliminate iron and sulfur, producing Bessemer matte in the case of nickel-copper smelters, and blister copper in the case of copper smelters

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Fig. 2 Preparation of third installment of unified computational framework for analysis of nickelcopper and copper smelters

it included basic mass and heat balances that were static and deterministic, and thus did not adequately incorporate concepts of bottleneck management. (As described by the Theory of Constraints [12], temporary variations surrounding a bottleneck, can effectively diminish the overall throughput of the entire system). Additionally, α was artificially fixed, hence a lack of thermochemical realism. The second instalment was presented at the second symposium on Pyrometallurgy of Nickel and Cobalt in 2017, and later published in Canadian Metallurgical Quarterly [10]. Much of the underlying concepts of the 2009 model were transferred into a discrete event simulation (DES) framework, which now supports extensible dynamics modeling, and random number generation. However, the DES framework of 2017 only applied a Gibbs Free Energy balance for the degree of oxidation α of the smelting furnace, and not for the converting aisle. The current paper is therefore to extend the DES framework, to adapt the approach of [10] which itself integrated earlier approaches of Goto [13], Kemori et al. [14], and Kyllo and Richards [15], and thus to incorporate dynamic α computations for each individual converter. Moreover, these developments involve dynamic temperature computations, which are related to the converter heat balance. This aspect is most notable to analyze the instalment of shrouded injectors and waterless matte granulation, thus to imbue PSC aisles with the advantages of SKS converters and Outotec-Kennecott flash converters, respectively. The shrouded injectors permit higher O2 enrichment in the converters, such that less N2 passes through the bath and into the offgas handling; the resulting offgas is thus richer in SO2, which is favorable for acid production. However, increased enrichment tends to overheat the converters, unless the blastrate is diminished (which actually decreases the overall throughput, ironically), or unless sufficient solid feeds are available to replace the cooling effect of N2. Indeed, the waterless matte granulation can be used to generate a stable supply of solid matte feed, and there is thus a synergistic benefit in simultaneously installing shrouded injection and waterless matter granulation. This approach for debottlenecking conventional smelters can allow conventional smelter designs to compete with both the SKS and Outotec-Kennecott designs. Nonetheless, smelters have been reluctant to experiment with this combination, perhaps because of a lack of computational tools to analyze the system-wide benefits.

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Both the 2009 cost model and the 2017 DES framework presented a system-wide perspective. However, the DES framework is extensible, meaning that it can be developed in phases, throughout an engineering project. Even in the early phases of the project, the framework presents a holistic representation of the smelter, with interconnected components that represents the main aspects of smelter dynamics. However, as the project progresses, the framework accommodates increasing levels of detail for each of the most critical aspects for a given project. Beyond the incorporation of dynamic mass and heat balances, the framework has already been used in the scope of complex industrial problems, e.g. to optimize production scheduling [16], and to manage environmental risk due to meteorological uncertainty [17]. Indeed, to execute smelter improvement projects, the DES framework provides a starting point to build a virtual smelter, and then to incorporate sufficient detail to quantify the hypothetical benefits and risks of proposed operational and technological changes. In summary, there are two avenues to extend the DES framework: (1) Incorporate features that are universally relevant to all nickel-copper and copper smelters (2) Customize the framework to address smelter-specific problems The current paper works toward the former, as the issue of dynamic thermochemical fluctuations within converter cycles is common to all conventional nickelcopper and copper smelters. Beyond the scope of the current paper, there is an interest to develop a computer library that will connect the framework to state-of-the-art thermochemical databases, thus to address deeper problems involving slag chemistry, including minor element transport. The third installment of the framework will therefore provide an invaluable connection between the system-wide smelter metrics, and the inner functioning of individual furnaces.

Dynamic Themochemical Modeling of Converter Cycles Peirce-Smith converters are operated in cycles, to produce batches of converted product [7, 8]. A cycle begins when an initially empty converter is filled to roughly half of its volume with matte, which is often combined with a quantity of silica flux as well as reverts, scrap and other cold charge. Oxygen-enriched air is then blown into the matte, as the iron combines with the incoming flux to form slag, and the corresponding sulfur is expelled as offgas; this process is known as the slag-blow, and it is halted intermittently, to skim away the accumulated slag, and possibly to introduce additional fresh ladles of matte and/or cold charges. Approximately all of the iron is eventually oxidized and skimmed away, thus signaling the end of the slag-blow. In the case of copper smelters, the blowing is continued, as the Cu2S (a.k.a. white metal, [8]) is converted into blister copper; this is known as the copperblow. Finally, the cycle is completed when the converter is emptied of its converted

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product, be it iron-free matte in the context of nickel-copper smelting, or blister copper in the context of copper smelting. Previously, the DES framework had considered only the discrete operational dynamics of converter cycles, including the initial charging and recharging events, as well as the skimming and emptying events. Indeed the durations of the blowing actions, as well as the quantities of incoming and outgoing material had all been treated as user parameters. The current developments are to explicitly represent thermochemical state variables PSC, to thus represent the evolving mass and heat content within each converter. This additional level of detail can allow the framework to evaluate (and eventually optimize) operating practices in response to short-term thermochemical variation. For example, if the temperature is rising slower than expected during a blowing action, then perhaps the skimming and recharging actions could be postposed, or perhaps certain types of cold charges could be conserved until a later charging event. Through simulation-based stochastic optimization [18], the framework can determine what operational triggers would favor one mitigating action over another. For each converter, the framework now considers the variables listed in Table 1, in which • • • •

nij is the number of moles of species or element i in phase j qBath is the heat content of the bath TBath is the common temperature of the slag and matte (i.e. TBath = TMatte = TSlag) α is the Fe3+/Fe2+ ratio in the slag, a.k.a. the degree of oxidation

The internally computed variables TBath and α, are not explicitly represented as state variables, but are part of intermediate computations to compute rate-changes, as described below. Additionally, the new formulation considers several operational parameters, • • • •

ϕ is the oxygen enrichment of the blast, expressed as a volume fraction ṅO2, Blast is the molar rate at which oxygen is blown into the bath rTarget is the target SiO2/Fe mass ratio within the slag TBlast is the temperature of the incoming blast

Table 1 Classification of converter thermochemical state variables Type of blow

Discretely dynamic state variables

Continuously dynamic state variables

Internally computed variables

Slag-blow in nickelcopper smelter Slag-blow in copper smelter Copper-blow

nNiS, Matte, nCoS, and nCu2S, Matte nCu2S, Matte

nFeS, Matte, nFe,Slag, nSiO2,Slag and qBath nFeS, Matte, nFe,Slag, nSiO2,Slag and qBath nCu2S, Matte, nCu, Blister and qBath

TBath and α

none

Matte

TBath and α TBath

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• TFlux is the temperature of the incoming flux • ϵ is the portion of incoming oxygen that reacts with the iron and sulfur rather, a. k.a. the oxygen efficiency • q̇Env is the rate of heat loss to the environment some of which can be altered by a simulated operational policy, depending on the nature of the simulation. Indeed, ṅO2, Blast can be computed from the blastrate and the oxygen enrichment ϕ. Within the scope of a simulation, these operational parameters are regarded as piecewise constant. In reality, certain of these quantities should actually be treated similarly to α and TBath, as internally computed state variables rather than operational parameters. In particular, ϵ and q̇Env should depend on the thermochemical and fluid mechanical conditions within the bath; in the context of a smelter-specific study that emphasizes oxygen utilization or heat transfer, the DES can indeed be extended to incorporate submodels that describe the dynamic behavior of ϵ and/or q̇Env . Otherwise, ϵ can be set to between 85% and 100%, and q̇Env can be set to between 2 and 4 MW, depending roughly on the configuration and size of the converter. Historically, discrete event simulation only considered state variables that are updated at discrete points in time, i.e. at discrete events. As a first approximation for nickel-copper PSC, nNiS, Matte, nCoS, Matte and nNiS, Matte are incremented only at charging events, and are decreased (to zero) only at the ends the cycles. More advanced implementations of DES allow the incorporation of state variables that have continuous dynamics, through the use of time-adaptive finite difference techniques, most notably the Runge-Kutta-Fehlberg method [19]; the incorporation of continuously changing variables into DES has been described the context of smelter dynamics by Navarra et al. [10]. The variables that are implemented with continuous dynamics (in addition to discrete dynamics) are more computationally demanding than those that only feature discrete dynamics. From the modeling perspective, however, the main difference is that the continuously changing variables require expressions for the rate of change. The following formulas focus on the slag-blow, assuming a fayalite slag, considering that a similar approach can also be applied for olivine slag. (The copperblow is far simpler than the slag-blow, because it is determined entirely from elemental mass balances that are algebraically decoupled from the heat balances). The rates of change of nFe,Slag, nFeS, Matte and nSiO2,Slag are determined through elemental mass balances that depend on α,   4 + 4α ṅFe, Slag = ϵṅO2, Blast 6 + 7α

ð1Þ

ṅFeS, Matte = − ṅFe, Slag

ð2Þ

( ṅSiO2, C1 =

 ṅFe, Slag , 0,

rTarget MFe MSiO2

r = rTarget r > rTarget

ð3Þ

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279

in which Mi is the molar mass of i, and r is the mass ratio of SiO2/Fe in the slag, as calculated from the aforementioned state variables r = (MSiO2nSiO2,Slag)/(MFenFe,Slag). In Eqs. 2 and 3, the dependency on α is made explicit by substituting for ṅFe, Slag . Equation 3 considers that r may exceed the target value rTarget following the addition of a cold charge, so that the continual addition of silica flux is halted until either r = rTarget is reestablished through continued slag-blowing, or until the next skimming operation, which ever comes first. The case where r < rTarget is not considered in Eq. 3, since it is presumed that the missing flux can be “rapidly” added, so that r = rTarget is instantaneously reestablished. However, in a real smelter, it may take some time before the operators realize that there is a shortage of flux. Therefore, in an engineering project in which the flux response is critical, a submodel can be developed that describes the operators’ behavior, including a probabilistic response time, and the risk of over- or under- compensating; otherwise, Eq. 3 is sufficiently detailed. The rate of change in qBath is dependent on both TBath and α. The heat-rate balance is described by q̇Bath = q̇Blast + q̇Flux − q̇Offgas − q̇Env ; in terms of molar flow rates, q̇Bath = ðṅN2, Blast ΔHN2, Blast + ṅO2, Blast ΔHO2, Blast Þ + ṅSiO2 ΔHSiO2, Flux   − ṅN2, Offgas ΔHN2, Offgas + ṅO2, Offgas ΔHO2, Offgas + ṅSO2, Offgas ΔHSO2, Offgas − q̇Env

ð4Þ in which the enthalpies of formation ΔHij of species i in stream j takes the general form,       bi  2 1 1 di  3 Tj − T o2 − ci Tj − T o3 ΔHij = ΔHio + ai Tj − T o + − o + Tj T 2 3

ð5Þ

in which in (ΔHjo , aj , bj , cj , dj ) are thermochemical constants [20], and T o = 298.15 K is the Standard Temperature. For simplicity, SiO2 has been assumed to be the only species within the flux and, conversely, flux is assumed to be the only continual stream of incoming SiO2, either through a chute or Garr Gun [7, 8]. Also, ṅN2, Blast = ṅN2, Offgas and ṅO2, Offgas can be computed from operational parameters ϕ, ṅO2, Blast and ϵ. Additional considerations are that the outgoing offgas has the same temperature as the bath, TOffgas = TBath

ð6Þ

and that the oxidation of FeS is such that ṅSO2, Offgas = ṅFe, Slag

ð7Þ

Within Eq. 4, the dependency on TBath is made explicit by applying Eqs. 5 and 6 to substitute for the ΔHi,Offgas factors. The dependence on α is made explicit by

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applying Eq. 3 to substitute for ṅSiO2 , Eq. 7 to substitute for ṅSO2, Offgas , and finally by applying Eq. 1 to substitute for ṅFe, Slag . Equations 1–7 give the rate-changes of continuous variables as a function of (TBath,α). However, (TBath,α) are themselves determined by a nonlinear system of two equations and two unknowns, as described in the Appendix. These values are solved by Newton Method [21], with the kth iteration given by 

TBathk αk



 =

TBath, k − 1 αk − 1

"

 −

∂fH ∂fG ∂TBath ∂α

1 −

∂fH ∂fG ∂α ∂TBath

∂fG ∂α G − ∂T∂fBath



∂fH ∂α

∂fH ∂TBath

#

fH fG

 ð8Þ

in which all instances of fH and fG, and their partial derivatives, are evaluated at (TBath,k-1,αk-1); typical values for (TBath0, α0) can be (1473 K, 0.15). As described in the Appendix, fH is proxy function, such that fH = 0 if and only if the heat balance is satisfied; likewise fG is a proxy function, such that fG = 0 if and only if the Gibbs Free Energy balance is satisfied. Equation 8 captures the coupled nature of TBath and α. The Newton iterations converge to a sufficient level of precision in usually fewer than ten iterations, and often fewer than five, giving values for TBath and α. These values are then used to update the rate-changes of Eqs. 1–7, which fit into a time-adaptive finite difference scheme to update the continuously dynamic state variables (Table 1), hence a dynamic representation of blowing actions. Intermittently, there are discrete instances in which material is charged or discharged from the converter. For simplicity, the skimmed slag is assumed to carry away the heat that corresponds to the bath temperature TBath, so that the skimming action does not alter TBath; however, within a more detailed smelter-specific project, it may be worthwhile to deduct an additional amount of heat that corresponds to the added radiation losses for having the hood lifted [7, 8]. In general, the formulation described in this section, and in the Appendix, provides a first approximation of PSC dynamics. Additional levels of detail are developed in phases, depending on the specific requirements of a particular project; this may involve modifying the equations described in this section, as well as developing submodels and subsubmodels. Ultimately, it is this extensible characteristic of DES that gives the framework its wide-ranging application.

Preliminary Computations Equations 1–8 have been implemented in Matlab® to provide preliminary results (Fig. 3), prior to implementation into the DES framework; the parameters are described in Table 2, in which wiFMatte is the weight fraction of i within the furnace matte. A total of 80 t of furnace matte is charged entirely at the beginning of the

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Fig. 3 Sample results displaying semi-discrete PSC dynamics

Table 2 Operational parameters Parameter

Value

ϕ ṅO2, Blast rTarget TBlast TFlux ϵ

23 360 0.7 298 298 85

vol.%O2 kmol/h K K %

Parameter

Value

q̇Env TFMatte wFeS, FMatte wNiS, FMatte wCoS, FMatte wCu2S, FMatte

2 1225 39 35 4 22

MW K % % % %

cycle, at a temperature of 1225 K, and is subject to a simplified slag-blowing operation. The operation is decomposed into five blowing actions, which are punctuated by vertical decreases in temperature that are triggered when the temperature reaches 1600 K. These decreases consist of skimming actions, which are immediately followed by the loading of cold charges that consist entirely of granulated Bessemer matte; sufficient cold charge is added so that the temperature falls to 1400 K. The repeated addition of these cold charges causes the bath to gain

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thermal inertia (heat capacity) as it accumulates nonferrous sulfides; this allows the blowing actions to become increasingly long, without overheating. The sample computations in this section describe an unorthodox converter cycle that could not be applied without assuring a steady supply of granulated Bessemer matte. Such a supply of cold charges could be provided by waterless matte granulation [4], for example. A similar approach could be adapted for copper smelters if, prior to the beginning of certain copper-blows, a portion of the white metal is granulated to be used later as cold charge. Moreover, the system-wide effect of such a strategy cannot be quantified unless dynamic thermochemical PSC computations are integrated within a broader framework.

Future Work Further development or extension of thermochemical capabilities of the DES framework can be achieved by using an add-in thermochemical library capable of carrying out complex multiphase Gibbs Energy Minimization for thermodynamic systems consisting of multiple components. This approach is particularly advantageous in the case of nonferrous pyrometallurgical processes, as these systems involve phases that are described by non-ideal solution models (see Appendix, or [10, 13–15]). More generally, a thermochemistry-based process simulation using Gibbs Energy Minimization reduces the number of assumptions, resulting in a very predictive and flexible tool. For example, the user does not need to explicitly specify the set of chemical reactions describing the process, nor the extent to which the reactions occur, nor the resulting phases. A complete thermochemistry-based process simulator requires three main components, namely: (1) a robust Gibbs Energy Minimization algorithm for multiphase multi-component systems, (2) solution models that describe the various non-ideal mixtures in a wide range of temperature and compositions and (3) an extensive and critically assessed thermodynamic database defining the required parameters for the system of interest. Currently, M4D-DES™ is being developed by M4 Dynamics Inc. as an Add-in for a DES framework (Arena™). M4D-DES is based on ChemApp™, a robust and complete Gibbs Energy Minimizer library compatible with an extensive number of solution models and compatible with all the FactSage-Family of thermodynamic databases as well as with user-created databases. This will provide increased applicability and flexibility to the DES framework described here. For example the extended framework will be able to analyze transport of minor elements, extend the number of system components of the system, and provide possibilities to better analyze the impact of different feed sources on the smelter throughput, including economic and environmental impacts.

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Appendix: Proxy Functions for Heat and Gibbs Free Energy Balances Equation 8 describes the Newton iterations that are used to simultaneously solve for the temperature of the bath TBath and the degree of oxidation of the slag α. These iterations are configured so as to simultaneously balance the heat and Gibbs Free Energy, hence two equations and two unknowns. Firstly, the heat balance is given by q = ∑ ∑ nij ΔHij in which S j is j ∈ fMatte, Slagg i ∈ S j

the set of species in bath stream j; equivalently the heat balance can be expressed in terms of a proxy function fH ≡ ∑ ∑ nij ΔHij − q, such that the fH = 0 if j ∈ fSlag, Matteg i ∈ S j

and only if the heat balance is satisfied. Following a sequence of substitutions using Eq. 5, and several additional algebraic manipulations, the proxy function can be expressed explicitly in terms of TBath and α, as required by Newton’s method.        2    1 1 1 fH ðTBath , αÞ = Σnij ΔHio + Σnij ai ðTBath − T o Þ + Σnij bi TBath − T o2 − Σnij ci − o 2 TBath T  3  1 o3 + Σnij di TBath − T 3 2 3 bFeO  2 o TBath − T o2 + aFeO ðTBath − T o Þ +   ΔHFeO 7 2 nFe, Slag 6 7 6   + 5 1+α 4 1 1 dFeO  3 o3 TBath − T − cFeO − o + TBath T 3  2  2 o o2 3   A + BðTBath − T Þ + C TBath − T α 6 7   + nFe, Slag 4  3 5−q 1 1 1+α − o + E TBath − T o3 +D TBath T

ð9Þ in which the summations are taken over (i, j) ∈ {(FeS, Matte), (NiS, Matte), (CoS, Matte), (Cu2S, Matte), (SiO2,Slag) }, and the constant coefficients (A, B, C, D, E) are given in Table 3. Given Eq. 9, it is straightforward to obtain expressions for ∂fH ∂fH ∂TMelt and ∂α , which are also present in Eq. 8. Navarra et al. [10] relied on a proxy function fG that was expressed explicitly as a function of α. The approach has now been slightly modified so that fG can be expressed explicitly as a function of TBath, as well as α.

Table 3 Coefficients for Eq. 9 A  1

o 2 ΔHFe3O4

o − ΔHFeO



B

C

1 2 ðaFe3O4

1 4 ðbFe3O4

− aFeO Þ

D − bFeO Þ



E 1 2 ðcFe3O4

− cFeO Þ

1 6 ðdFe3O4

− dFeO Þ

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̸ TBath

fG ðTBath , αÞ = ∏ ðAl + Bl αÞCl + Dl

9

− ∏ ðAl + Bl αÞCl + Dl

l=1

̸ TBath

ð10Þ

l=4

in which the coefficients (Al, Bl, Cl, Dl) are given in Table 4. Considering that Dl is zero for all but two factors, Eq. 10 can also be expressed as 2

fG ðTBath , αÞ = ∏ ðAl + Bl αÞCl ðA3 + B3 αÞ15430

̸TBath

8

− ∏ ðAl + Bl αÞCl ðA9 + B9 αÞ15430

l=1

̸TBath

l=4

H which is helpful for finding an expression for ∂T∂fMelt ; to find an expression for ∂f∂αG it is simpler to work directly with Eq. 10. Within Table 4, R denotes the ideal gas constant, and XFeS, Matte is the molar fraction of FeS within the matte, given by XFeS, Matte = nFeS,Matte /(nFeS,Matte + nNiS, Matte + nCoS, Matte + nCu2S, Matte). Additionally, Table 4 includes ΔH0 and ΔS0 which are the enthalpy and entropy of the following equilibrium,

FeSðMatteÞ + 3Fe3 O4ðSlagÞ ↔ 10FeOðSlagÞ + SO2ðOffgasÞ Indeed, the approach of Navarra et al. [10], and hence Eq. 10, is based on this equilibrium, using the expressions of Goto [13] and Kemori et al. [14] for the activity coefficients of FeS, FeO and Fe3O4. These expressions had originally been developed for copper smelters, but were successfully adapted and tested for nickelcopper smelters [15].

Table 4 Coefficients for Eq. 10 l

Al

Bl

Cl

Dl

1 2 3

1 2 

1 -1

1 10 0

0 0 15430

4

XFeS, Matte e 0  

0

1

0

1 

3 1

0 0





3

0

4

0

0

15430

5 6 7 8 a

a

9

2.44 − 0.4



nSiO2, Slag nFe, Slag



ΔS0 R

3 − ϕϵ 2ϕϵ

  n Slag 1.38 + 12.28 nSiO2, Fe, Slag    n Slag 2 1 + nSiO2, Fe, Slag    n Slag K 2 1 + nSiO2, Fe, Slag

in which K = e − ΔH0

̸15430R

   n Slag − 1.42 + 0.4 nSiO2, Fe, Slag



7 − 3ϕϵ 4ϕϵ



56.8 + 12.28   2 2



nSiO2, Matte nFe, Slag nSiO2, Slag nFe, Slag

 K

nSiO2, Slag nFe, Slag



ð0.54 + 0.52XFeS, Matte + 1.4XFeS, Matte ln XFeS, Matte Þ1458

̸15430

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References 1. Price T, Harris C, Hills S, Boyd W, Wraith A (2009) Peirce-smith converting: another 100 years? In: Paper presented at the 138th TMS annual meeting, San Francisco, California, 15–19 February 2009 2. McRae M (2018) Nickel. Minerals Commodity Summary. United States Geological Survey, Washington DC, pp 112–113 3. Flanagan D (2018) Copper. Minerals Commodity Summary. United States Geological Survey, Washington DC, pp 52–53 4. Mucciardi F, Palumbo E, Jin N (1999) A waterless caster for matte/slag granulation. In: Paper presented at the 4th international copper conference, Phoenix, Arizona 10–13 October 1999 5. Navarra A, Kapusta J (2009) Decision-making software development for incremental improvement of nickel matte conversion. In: Paper presented at the 48th annual conference of metallurgists of CIM, Sudbury, Ontario, 23–26 August 2009 6. Chibwe D, Akdogan G, Bezuidenhout G, Kapusta J, Bradshaw S, Eksteen J (2015) Sonic injection into a PGM Peirce-Smith converter: CFD modelling and industrial trials. J S Afr I Min Metall 115:115–349 7. Crundwell F, Moats M, Ramachandran V, Robinson T, Davenport W (2011) Converting— final oxidation of iron from molten matte. In: Chapter 19 in: Extractive metallurgy of nickel, cobalt and platinum group metals. Elsevier, Oxford, pp 233–246 8. Schlesinger M, King M, Sole K, Davenport W (2011) Converting of copper matte. In: Chapter 8 in: extractive metallurgy of copper. Elsevier, Oxford, p. 127–153 9. Ramirez C, Ruz P, Riveros G, Warczok A, Treimer R (2009) Chrome-magnesite refractory corrosion with olivine slag of high cuprous oxide content. In: Paper presented at the 138th TMS annual meeting, San Francisco, California, 15–19 February 2009 10. Navarra A, Valenzuela F, Cruz R, Arrancibia C, Yañez R, Acuña C (2018) Incorporation of matte-slag thermochemistry into sulphide smelter discrete event simulation. CMQ 57(1):70–79 11. Ng K, Kapusta J, Harris R, Wraith A, Parra R (2005) Modeling Peirce-Smith converter operating costs. JOM 57(7):52–57 12. Dettmer H (2007) Introduction to the theory of constraints. In: Chapter 1 in: the logical thinking process: a systems approach to complex problem solving. ASQ Quality Press, Milwaukee, p. 3–30 13. Goto S (1974) Equilibrium calculations between matte, slag and gaseous phases in copper smelting. In: Paper presented at the annual meeting of the institution of mining and metallurgy, Brussels, Belgium, 11–13 February 1974 14. Kemori N, Kimura T, Mori Y, Goto S (1987) An application of Goto’s model to a copper flash smelting furnace. In: Presented at the annual meeting of the institution of mining and metallurgy, London, England, 21–23 September 1987 15. Kyllo A, Richards G (1991) A mathematical model of the nickel converter: part 1—model development and verification. Metall Trans B 22(2):153–161 16. Navarra A (2016) Automated scheduling and scientific management of copper smelters. T I Min Metall C 125(1):39–44 17. Navarra A, Marambio H, Oyarzún F, Parra R, Mucciardi R (2016) System dynamics and discrete event simulation of copper smelters. Miner Metall Process 34(2):96–106 18. Kleijnen J, Wan J (2007) Optimization of simulated systems: OptQuest and alternatives. Simul Model Pract Th 15(3):354–362 19. Kelton W, Sadowski R, Swets N (2010) Continuous and combined discrete/continuous models. In: Chapter 11 in: Modeling and simulation of discrete-event systems. Wiley, Hoboken, New Jersey, pp 473–512 20. Liley P, Thomson G, Daubert T, Buck E (1997) Physical and chemical data. Part 2 in: Perry’s Chemical Engineers’ Handbook. McGraw-Hill, New York, pp 1–204 21. Burden R, Faires J, Burden A (2016) Numerical solutions of nonlinear systems of equations. In: Chapter 10 in: numerical analysis. Cengage Learning, Boston, Massachusetts, pp 641–683

Development of Continuous Radar Level Measurement for Improved Furnace Feed Control Rodney Hundermark, Quintin van Rooyen, Paul van Manen, Chris Steyn, Afshin Sadri and David Chataway

Abstract At Anglo American Platinum’s Polokwane Smelter, radar instruments are being used to provide continuous measurement of the concentrate feed level in the six-in-line electric furnace used for smelting of nickel-copper concentrates containing platinum group metals (PGMs). The radar instruments are installed directly on the furnace roof, and a mechanical system has been designed to protect them from radiation and elevated freeboard temperatures and pressures. Signal validation criteria have been implemented to ensure that proper level measurements are being used by the PLC system. The feedback from the radars has provided an improved understanding of the furnace behaviour with respect to concentrate feed level control, individual feeding events, global feed rates, feed distribution and bath disturbances. In combination with a dynamic mass balance for continuous calculation of liquid level in the furnace, the radar measurements have enabled more precise feedback control of the concentrate blacktop depth. In turn this has yielded improvements in furnace stability. Keywords Electric furnace



Level measurement



Radar



Control

R. Hundermark (✉) Anglo American, 45 Main Street, Johannesburg, South Africa e-mail: [email protected] Q. van Rooyen ⋅ P. van Manen Polokwane Smelter, R37, Polokwane, South Africa C. Steyn Anglo American Platinum, 55 Marshall Street, Johannesburg, South Africa A. Sadri ⋅ D. Chataway Hatch, 2800 Speakman Drive, Mississauga, Canada © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_23

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Introduction Anglo American Platinum’s Polokwane Smelter is located approximately 10 km south of the city of Polokwane in the Limpopo Province in South Africa. The smelter processes nickel-copper concentrates containing platinum group metals (PGMs) to produce an enriched furnace matte which is further processed elsewhere. The main asset at the smelter is a six-in-line electrode, rectangular, 68 MW electric furnace in which the concentrate is smelted. Details of the operation have been provided previously [1, 2]. An important control variable on any electric furnace is the power-to-feed ratio, or specific energy consumption (SEC in MWh/t). Overpowering of a furnace with insufficient feed can lead to catastrophic failure of the furnace lining within hours in the worst case. In the PGM smelting industry, the power-to-feed ratio in the primary electric smelting furnaces can be crudely, but robustly, managed through control of the thickness of the unmelted concentrate layer floating on top of the slag (referred to as blacktop thickness). That is, if the furnace is underfed then the blacktop thickness will be low, and if the furnace is overfed, then the blacktop thickness will be high. Measurement of the blacktop thickness, and matte and slag levels is a manual operation requiring operators to go onto platforms above the furnace roof. Normally the furnaces are operated under a slight negative pressure, however, it is possible for positive pressure events to occur, in which hot gas and dust may be ejected from the furnace. From a hierarchy of controls perspective, efforts are being made by Anglo American to eliminate or reduce the exposure of operating personnel to the molten materials environment, and also provide improved monitoring and control. Such efforts have included the implementation of new technologies such as fibre optic temperature instrumentation [3], advanced statistical event detection tools [4] and improved instruments for concentrate feed level measurement, such as radars. Testing of radars for concentrate feed level measurement in the electric furnace commenced as far back as 2009, and ultimately led to a full industrial installation of eight radars on the Polokwane furnace during 2013. The main objectives for the full-scale installation were improved safety considerations (in decreasing the time operators need to spend above the furnace roof level) and improved furnace power-feed ratio control leading to improved furnace efficiency. The purpose of this paper is to discuss the radar performance, and how the measurements have been utilised for improvement of the stability and narrowing the operating range of furnace level control.

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Background on Radar and Deployment The existing technique for determining furnace levels is a manual sounding technique. These measurements are made by means of a sounding bar, which is a ∼6 m long steel rod lowered down into the furnace with a winch. Slag and matte freeze onto the bar, and a light dusting of concentrate from the concentrate bed adheres to the bar (Fig. 1). The sounding bar is raised back above the furnace roof level and the operator makes use of a graduated measuring bar which is held up adjacent to the sounding bar, and the levels of matte, slag and blacktop are determined. This technique is relatively simple in principle, but somewhat prone to inaccuracies due to difficulties to distinguish accurately the top of concentrate feed and matte-slag interface level. Moreover, soundings expose operators to risks above the furnace roof and can only be taken relatively infrequently, typically at most every hour [6]. In contrast the radar instruments can only measure the top of the concentrate feed level (Fig. 1), however, they provide continuous feedback of this level. It needs to be emphasised that the radar feedback needs to be combined with the liquid level information (build-up + matte + slag) inside the furnace obtained from the manual soundings in order to yield a full understanding of the furnace levels. Much of the background information on the radar technologies employed at Polokwane Smelter has been published elsewhere [5, 6] in 2011 and 2017, respectively. Radars operate by emitting and receiving electromagnetic waves, which are then reflected by interfaces of conductive materials or materials with

Fig. 1 Illustration of the sounding (left) and radar (right) level measurement techniques in an electric furnace

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different dielectric properties. Most radars are either pulse-based or continuouswave frequency-modulated (FMCW). Pulse-based radars measure the time required for a transmitted microwave pulse and its reflected echo to make a completed return trip between the non-contacting transducer and the measured object’s surface [7]. FMCW radars instead change their operating frequency during the measurement and measure the difference in phase or frequency between the transmitted and received microwaves [8]. With either technique, the transceiver in the radar converts the received signal electrically into a distance/level reading and outputs it in either an analogue and/or digital form [7]. Radars have been commonly used to measure the level of liquids and solids in vessels, such as oil storage tanks [9] and grain silos [10] but have been difficult to install in harsh conditions such as extreme heat and electromagnetic interference (EMI), as seen in electric furnaces. Electric furnace freeboard temperatures can be higher than 1000 °C and the electrodes inside carry kiloamperes of current, creating significant EMI. Radar-based measurement systems are highly effective for overcoming the challenges present in the furnace freeboard: dust and vapour, which would be highly limiting for a visual/laser based system, are all but invisible to radar given the appropriate selection of its operating frequency. Furthermore, temperature fluctuations during furnace operation, which would be highly limiting for an acoustic based system, have little effect on the accuracy of level measurements (i.e. 0.026% difference from 0 °C to 2000 °C [11]). To combat heat and EMI, Hatch has designed an enclosure and monitoring system to protect the radar units during operation. The enclosure is metal, acting as a Faraday cage which blocks out EMI and increases the signal to noise ratio. Supplied cooling air keeps the enclosure within the radar’s acceptable operating temperature range, which is verified by temperature monitoring. The supplied air also protects the antenna from dust. A carefully chosen refractory cloth blocks thermal radiation from the freeboard while permitting the transmission of microwave radiation and further limiting the exposure of the radar antenna to dust. To date, the barrier cloth has proven to work well, demonstrating the ability to withstand the rigours of the furnace freeboard environment. While the barrier cloth is well suited to protect the radar, it is still exposed to extreme conditions and as a result, may need to be replaced infrequently to maintain mechanical integrity. In addition, the barrier cloth allows the radar signal to penetrate through the barrier with transmission coefficients of greater than 95%. The design and components of the cooling system (the thermal barrier, compressed instrument air inlet and diagnostic instruments) were optimized based on analysis and thermal modelling (Fig. 2). Radar units from six different vendors were investigated. From a shortlist of six different radar units, three units were selected and each were successfully installed and tested at Polokwane within a few months. The specifications and differences between those units tested are highlighted in Table 1. The primary selection criterion was the radar’s maximum operating temperature as it was foreseen that irregular periods of extremely high temperature could be experienced, even with a cooling system installed. Prior to the selection process, both types of radars measurement technique (pulsed and FMCW) were studied in

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Faraday cage

Barrier cloth

Fig. 2 Radar enclosure installed on the Polokwane roof (left) and typical thermal modelling of a radar with a cooling system on a furnace roof (right)

Table 1 Select radar specifications Category

Unit A

Unit B

Unit C

Radar operating principle Frequency (Bandwidth) Max. operating temperature Polarization

Pulse 26 GHz 80 °C Circular

FMCW 24.2 GHz (1 GHz) 65 °C Linear

FMCW 78 GHz (1 GHz) 80 °C Linear

detail and various units were tested in the laboratory and it was determined that both techniques seemed feasible for implementation on a furnace. In addition, the effect of radar polarization was investigated: radar microwaves are generated either in elliptical (including circular) or linear polarization format according to the geometrical orientation of the field oscillations generated. In the studies it was found that the radars which emit circularly polarized microwaves were more advantageous for the furnace application in comparison to radars that emitted linearly polarized microwaves. Circular polarization of microwaves can limit the risk of false echoes from unknown sources and can allow the device to distinguish between echoes that have been reflected by one surface from those that have been reflected twice [6]. This allows for the unit to more easily identify the reflection from the charge material and to be placed more freely on the furnace roof considering the proximity to vertical surfaces such as electrodes or walls. Eight radars were installed on the Polokwane furnace in 2013 with the intent to provide spatial coverage over a large area of the furnace (Fig. 3). The objective was to observe the difference in the blacktop thickness over the length of the furnace as a function of the quantity of feed provided at the two feed ports adjacent to each radar unit. It was anticipated that individual feeding events would be detectable by the radar when the local blacktop thickness increased.

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R. Hundermark et al. Radar E

Port W3

Radar F

Port W4

Radar H

Radar G

Slag end

Matte end

Radar D Feed port

Radar A

Radar B

Radar C Radar port

Fig. 3 Schematic of feed port and radar port layout relative to electrodes

Validation of Radar Data

350 300 250 200 150 100 50

Average build-up sounding

Average matte sounding

Average slag sounding

Average blacktop sounding

Average radar

Hourly moving average feed rate

Furnace power

Fig. 4 Trend of average matte, slag and blacktop levels for a 3 day period

06-Mar 05:00

06-Mar 02:00

05-Mar 23:00

05-Mar 17:00

05-Mar 20:00

05-Mar 14:00

05-Mar 11:00

05-Mar 05:00

05-Mar 08:00

05-Mar 02:00

04-Mar 23:00

04-Mar 20:00

04-Mar 17:00

04-Mar 14:00

04-Mar 11:00

04-Mar 05:00

04-Mar 08:00

04-Mar 02:00

03-Mar 23:00

03-Mar 20:00

03-Mar 17:00

03-Mar 14:00

03-Mar 11:00

03-Mar 08:00

0 03-Mar 05:00

Average soundings, radar level (cm), feed rate (t/h) and furnace power (MW)

One of the most important considerations when commissioning the radars was validation of the measurements. The validation was carried out by comparing the radar measurements to manual measurements of the top of concentrate level (Fig. 1). Soundings on the Polokwane furnace are typically measured at 2 or 3 ports every hour. The level data is captured manually into the historian, while the radar information is reported back to the PLC/ SCADA/ historian continuously. In order to illustrate the types of measurements obtained from the soundings and the radars, the average matte, slag and blacktop levels are provided for a 3 day period (Fig. 4).

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As the sounding bar measurements are discrete events, there are steps in the average levels, while the radar provides continuous feedback. Inspection of the trends (Fig. 4) yields the following observations: • The furnace entered a shutdown period (pink line steps down to zero MW on 3 Mar at 18:00) with a corresponding decrease in the hourly moving average feed rate (green line). • Ahead of this time blacktop thickness was decreased intentionally by reducing the feed rate as indicated by both the sounding information, and the radar feedback (blue line). • On restart of the furnace after ∼24 h, the radars indicated that the blacktop thickness dropped to the top of slag level, as would be expected for introducing power into the furnace with no feed. • With commencement of feeding at around 01:00 on 5 Mar there was an almost instantaneous increase in the radar blacktop measurement, which was followed by an increase in the sounding blacktop thickness measurements. • It can be observed that there is not perfect agreement between the sounding blacktop thickness when compared to the radar measurements. When plotting the radar blacktop thickness measurements versus the sounding measurements (Fig. 5) there is reasonable agreement between the data when there are significant changes in the blacktop thickness, e.g. when the furnace is powered down and the concentrate levels are decreased, however, there is greater spread at the normal operating blacktop thickness levels (80–110 cm, as measured by radar). Possible reasons for this are discussed further below. The above analysis provides a global view of the furnace feeding and level control and provided a degree of confidence that the radar measurements could be utilised for control purposes. However, given the number of radars installed it was

160

Fig. 5 Radar blacktop thickness versus sounding blacktop thickness measurements Radar blacktop thickness (cm)

140

120

100

80

y = 0.8602x R² = 0.8303

60

40

20

0

0

20

40

60

80

100

120

Sounding blacktop thickness (cm)

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anticipated that a more accurate local view of the furnace would be obtained. That is, feeding at local ports should contribute to local level increases at the radar instrument adjacent to the feed ports. This phenomenon was explored by means of evaluating normal feeding events over short time intervals, and then by means of step tests where feeding at specific feed ports was stopped and the response observed. An assessment was carried out of feeding at two feed ports adjacent to a radar instrument, as highlighted in the red rectangle in Fig. 3, over a two hour time period. The feedback from the radar instrument did not correlate well with the feeding at the two ports (Fig. 6) and many other disturbances in the radar measurement were apparent in times when feeding at the local ports was not in progress. The step tests were carried out on the same feed ports and radar as mentioned above. Feed was stopped at the two ports adjacent to Radar F for a period of two hours, and it was expected that there would be a relative decrease in the level measured at Radar F relative to the surrounding radar units (C to G). Based on the trends of the radar levels over the test duration (Fig. 7) there was no clearly discernible drop in the Radar F level (green line relative to the other lines). A hypothesis for this observation is that the dry concentrate which is fed into the furnace becomes fluidised when aerated, and the dynamic angle of repose of the concentrate is very low. It is postulated that the concentrate flows out and away from the local feed ports significantly further than originally envisaged, and therefore influences the radar measurements at locations much further away. Calculations were carried out on the influence of the angle of repose on an ideal feed pile height and radius

300

0.7

295 0.6

0.5

285 280

0.4

275 0.3

270 265

0.2

260 0.1 255 250

c410_LI_775F_Qlevel

Port W3 feed

Port W4 feed

Fig. 6 Feeding at local ports and response of adjacent radar (top of concentrate level)

07-Sep 08:10

07-Sep 08:00

07-Sep 07:50

07-Sep 07:40

07-Sep 07:30

07-Sep 07:20

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0

Feed rate (t/min)

Radar levels (cm)

290

Development of Continuous Radar Level Measurement …

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140

Blacktop thickness (cm)

120 100 80 60 40 20

Radar C blacktop thickness

Radar D blacktop thickness

Radar E blacktop thickness

Radar F blacktop thickness

Radar G blacktop thickness

Test start

04-Sep 18:00

04-Sep 17:00

04-Sep 16:00

04-Sep 15:00

04-Sep 14:00

04-Sep 13:00

04-Sep 12:00

04-Sep 11:00

0

Test stop

Fig. 7 Comparison of radar output during step test at Radar F

assuming a 3t feeding event of concentrate (typical batch size) with a bulk density of 1.5 t/m3 and a right circular cone pile geometry. For angles of repose of 6°, the concentrate will spread out to an approximate 2.5 m radius around the feed port, and result in a pile height increase of around 27 cm (Fig. 8). Given that the distances between feed ports range from 3.4 to 4 m on the Polokwane furnace, it is conceivable that feeding at one port will impact on the closest radar and others. When feeding normally at all feed ports (Fig. 6) the above calculations suggest that feeding on the eastern side of the furnace will influence the Radar F measurement. It is postulated that the low angle of repose for the concentrate explains why global, and not local, blacktop thickness measurement changes are observed. As noted above there are discrepancies between the sounding blacktop thickness measurements and the radar blacktop thickness measurements, particularly at the “normal” operating range for the blacktop thickness. Potential causes for this are the following: • There is a range of 25–30 cm in the feedback of the individual radar instruments (observable in Fig. 7). The spread tends to remain quite consistent, however, it appears not to be related to the relative elevation of the individual radars, i.e. they are all installed at the same height above the furnace roof. • The blacktop thickness measurement off the sounding bar relies on operator observation of the light dusting of the grey concentrate powder against the sounding bar. There is likely a +/- 5 cm range for the measurement.

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• The sounding blacktop thickness measurements are discrete, however, the radar measurements are continuous. The sounding data is inputted manually into the data historian and relies on a manually entered time and date stamp. Some offset in the time of the sounding data may be introduced in this manner. Regarding whether the radar measurements are of sufficient quality for control purposes, it is argued that they are for the following reasons: • The most significant risks to the furnace operation are those related to the furnace being underfed (risk of runout) and overfed (risk of significant positive pressure events when process gases escape during bath disturbances). Based on many instances of trends where the blacktop thickness has been decreased to near zero intentionally (e.g. during furnace shutdowns), the radar measurements have reliably reflected the significant drop in the concentrate level. Similarly, overfed conditions have also been detected. • The dynamics in the furnace are relatively slow, therefore the average change in blacktop thickness is relatively slow. By way of example, the cross sectional area of the Polokwane furnace is approximately 300 m2, and with a smelting rate of 90t/h and a bulk density of concentrate of 1.5 t/m3, the average change in concentrate level with no incoming feed would be 20 cm. Given a typical blacktop thickness of 80 cm, there is a 4 h residence time within the blacktop. This obviously provides an average view of the furnace, and in reality the concentrate level around the electrodes will decrease much more rapidly due to higher local temperatures. However, it is illustrated that the radars are not

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required to provide very short interval control for small blacktop thickness changes. A brief application of the basic sampling theorem is presented during the control strategy discussion. On this basis, further work was carried out on the implementation of improved furnace feed control to utilise the continuous radar measurements.

Development of Dynamic Mass Balance Models with Radar Input As noted above the manual sounding measurements of the matte, slag and blacktop thicknesses are carried out on an hourly basis. The measurements are therefore discrete in contrast to the continuous radar measurements. As the radars can only measure the top of concentrate level, knowledge is required of the liquid level (build-up + matte + slag) inside the furnace, and hence the discrete sounding measurements are required. However, the discrete nature of the sounding measurements gives rise to step changes in the radar blacktop thickness calculations. From a control perspective this is undesirable as the step changes introduce step changes into the control response and give rise to instability. To overcome this, an approach was developed for calculation of a dynamic mass balance for prediction of continuous matte and slag levels, such that a continuous blacktop thickness measurement could be derived: • The continuous furnace feed rate is known, as is the furnace power input, and therefore based on assumption of the specific energy consumption for a given concentrate composition and prevailing furnace heat losses, an estimated smelting rate can be calculated. • Based on the feed composition, the matte and slag generation rates can be calculated. • The quantity of matte tapped from the furnace is measured by means of load cells on the matte ladle transfer crane. • The quantity of slag tapped from the furnace is measured by means of a weightometer on the granulated slag conveyor and a correction is applied for the moisture content in the wet granulated slag. • The offgas dust losses are assumed to be in steady state, as the dust is returned to the furnace on a continuous basis, and are therefore not taken into account. Based on the above, the expected change in the top of blacktop thickness can be calculated on a continuous basis. Given that there are several assumptions contained in the above approach, there is risk that the dynamic model can diverge, and, to prevent this, regular “calibration” of the model is carried out using the sounding information for the various levels. To provide an indication of the output of the dynamic mass balance model, the predicted and sounding levels are illustrated for a 24 h period (Fig. 9). Of specific note are periods during which no slag tapping

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Fig. 9 Example of dynamic mass balance model output

occurs (pink line) and the model predicts the increase in the slag level on a continuous basis. The increase in level is then validated by the sounding measurement for the slag level. By means of this approach, a continuous bath liquid measurement could be provided for control purposes.

Development of Control Strategies with Radar Input Work continued to determine whether a possible furnace efficiency improvement could be achieved by including the radar measurement into the control algorithm. The current application of the discrete sounding measurements provides feedback to the Hatch Supervisory controller. Therefore, it was questioned whether a continuous feedback signal of the blacktop thickness could improve furnace stability and ultimately improve matching the concentrate fed relative to the applied power. The design of any control system is preceded by an investigation into the dynamic response and frequency signature of the process. A simple finite impulse response (FIR) model as presented by Ljung [12] describes the dynamic step response model of the blacktop thickness to a 1t/h change in concentrate feed setpoint (SP) in Fig. 10. The blacktop thickness in this instance is defined as the difference between the top of concentrate radar measurement and the modelled slag/concentrate interface level. The Laplace domain approximation of the blacktop thickness step response model in Fig. 10 is given as:

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Gs =

0.0000717e − 900θ 1000s2 + s

Some of the more notable features of this model are: • The integrating response of the blacktop thickness is similar to that of a liquid tank level with a constant flow out and an increase in incoming flow. • 900 s dead-time from the step in feed SP to the start of movement from the blacktop thickness • Steady state integration of 7.17 × 105 cm/s which is very similar to the rudimentary model described previously as 6.17 × 105 cm/s for every t/h (20 cm/h decrease observed for a 90 t/h cut in feed). • A lag of 1000 s is observed in the system, which, together with the dead-time indicates that the system reaches the steady-state ramp at approximately 2500 s. In a system perceived to be as slow moving as the furnace it might be argued that dynamic information at higher frequencies is not required. An analysis of the frequency response of the blacktop thickness to the concentrate feed is performed to invalidate this argument. To assist in understanding why the continuous measurement is more beneficial than the traditional once-an-hour discrete blacktop thickness sounding, the basic sampling theorem is considered [13]. The theorem states that: in order to observe all the dynamic information in a signal with dynamic components up to a maximum frequency of ωmax, the sampling frequency ωs must be at least twice the rate. ωs > 2ωmax

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Similarly, for a signal with a specific sampling frequency ωs it holds true that dynamic components are only observed at frequencies < ωs/2. In the case of the furnace sounding which is performed at a frequency of 2.7x104Hz (once an hour), dynamic information for the furnace is only available at frequencies of 1.35x104 Hz or lower, or periods > 2 h. However, when considering that the furnace blacktop thickness starts showing dynamic behaviour after 20 min and reaches a steady-state of integration after 42 min (Fig. 10) it can be argued that the sampling frequency should ideally be greater than once every 10 min to capture all dynamic components. Note that this frequency response is only for a change in feed SP, other dynamic component from various disturbances will also be present and might be quicker. Regardless, this finding presents a strong argument for the implementation of a continuous blacktop thickness feedback to the Hatch Supervisory algorithm. In order to use the blacktop thickness feedback from the 8 radar units as input to the furnace feed controller a minimum number of available radar units was defined to ensure accurate and reliable input. Thus the availability of these units became a key parameter in the successful integration and operation of the radar based feed controller. As noted above, the radar units are protected from the furnace freeboard dust and gas by a physical barrier and pressurization system. With longer term operation of the radars it was observed that the number of available units decreased over time, due to ingress of dust and gas into the instrumentation compartment caused by mechanical failure of the refractory cloth. Due to the location of the radar units they can only be accessed when the furnace is electrically isolated, therefore once the number of available units reaches the minimum threshold they can only be restored during a shutdown. Alternative barriers have been tested to provide protection against furnace freeboard conditions, and although radar availability has improved it still remains an area of focus. The blacktop thickness feedback in the Hatch Supervisory controller was changed from the manual soundings to a quasi-continuous measurement in the form of the difference between the radar’s top of concentrate and the sounding top of liquid level measurements. Commissioning of this change commenced during the first quarter of 2015. The decision to persist with the sounding liquid level was made as more test work and model robustness studies were required before the mass balance derived liquid level could be implemented. An improvement in the accuracy of the blacktop thickness control was achieved with a reduced standard deviation from an average of 37% relative standard deviation before the radar input, to an average of 24% relative standard deviation with the use of radar (Fig. 11). This has allowed for a reduction in the blacktop thickness setpoint to ∼80 cm which has helped lead to more stable operating conditions.

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Fig. 11 Blacktop thickness history and relative standard deviation for time periods pre- and post-implementation of radar control

Discussion and Conclusion The use of radar measurements has allowed for the first non-thermal continuous measurement of process parameters within the Polokwane furnace. Reasonable agreement was obtained between the radar measurements and the traditional blacktop thickness measurements using sounding bars, but importantly, large magnitude changes in the blacktop thickness were reliably detected by the radars. The continuous radar feedback was combined with the discrete liquid level measurements to provide the blacktop thickness on a continuous basis, however, the discrete changes in liquid level introduced step changes in the blacktop thickness. An improved continuous liquid level predictor was developed using a dynamic mass balance and provided insight into the transient level responses of within the furnace. When combined with a continuous liquid level prediction, the radar measurements provide for continuous feedback of the blacktop thickness—a key control parameter. With use of the radars, an improved feed control strategy has been implemented and has started to provide indications of improved blacktop thickness control with a better adherence to setpoints and a tighter standard deviation. Further improvements to the control strategy are currently in progress. These improvements are all aimed at improved furnace reliability, efficiency and longer term further automation of the furnace, with reduction in exposure of operating personnel to the molten materials environment.

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Acknowledgements The authors express their gratitude to Anglo American Platinum for permission to publish this paper.

References 1. Hundermark R, de Villiers B, Ndlovu J (2006) Process description and short history of Polokwane Smelter, Southern African Pyrometallurgy 2006. In: Jones RT (ed) South African institute of mining and metallurgy, Johannesburg, 5–8 March 2006, pp 35–41 2. Hundermark RJ, Mncwango SB, de Villiers LPvS, Nelson LR (2011) The smelting operations of Anglo American’s platinum business: an update, Southern African Pyrometallurgy 2011. In: Jones RT, den Hoed P (ed) Southern African institute of mining and metallurgy, Johannesburg, 6–9 March 2011 3. Sakaran RL, van Rooyen Q, van Manen PK, Mukumbe PP (2017) Analysis and interpretation of fibre optic temperature data at the Polokwane Smelter. In: 7th international platinum conference ‘Platinum—A Changing Industry’ in association with AMI precious metals 2017, 18–19 October 2017 4. Groenewald JWD, Nelson LR, Hundermark RJ, Phage K, Sakaran RL, van Rooyen Q, Cizek A (2017) Furnace integrity monitoring using principal component analysis: an industrial case study. In: 7th international platinum conference ‘Platinum—A Changing Industry’ in association with AMI precious metals 2017, 18–19 October 2017 5. Shameli E, Venditti R, Uyeda B, Kepes A, Gerritsen T, Sadri A, Southall S (2011) Enhanced furnace feed control using radar level measurement. In: European metallurgical conference proceedings 2011, pp 1607–1616 6. Braun W, Chataway D, Janzen J, Kepes A (2017) Feed and charge level measurements in furnaces. In: European metallurgical conference proceedings 2017, pp 679–694 7. Indumart, Various Technics of Liquids and Solids Level Measurements (Part 3). https://www. indumart.com/Level-measurement-3.pdf 8. Radartutorial.eu, Radar Basics Frequency-Modulated Continuous-Wave Radar (FMCW Radar). http://www.radartutorial.eu/02.basics/Frequency%20Modulated%20Continuous% 20Wave%20Radar.en.html 9. Emerson, About Radar Level for Tank Gauging. http://www.emerson.com/en-us/automation/ measurement-instrumentation/tank-gauging-system/about-radar-level-measurement-for-tankgauging 10. Vega, Grain Silo. https://www.vega.com/en/home_sv/Applications/Food-industry/Grain-silo 11. Devine P (2000) Radar level measurement—The user’s guide, VEGA Controls 12. Ljung L (1999) System identification: theory for the user. Second Edition, Prentice Hall PTR 13. Luyben WL (1990) Process modelling, simulation, and control for chemical engineers. 2th edn. McGraw-Hill Publishing Company

Research on Recovery of Valuable Metals in Waste Acid from Copper Smelting Flue Gas Acid-Making and Reduction and Harmless Treatment of Solid Wastes Yan Wen, Zhen Bao and Xinmin Wu

Abstract For the impurities in the copper smelting waste acid, SO3, Cu, As, Pb, Zn, Re and other elements enter the waste acid and generate a lot of arsenic residue, gypsum and neutralized residue, in which Cu, Zn and Re metals have not been recycled. Cu, Re and As are separated by step-wise vulcanization. The valuable metals of Cu, Zn and Re are recovered by separating zinc through pH adjustment. Calcium injection is conducted for boiler for reducing the amount of SO3. The residues generated in the form of arsenate and hydroxide will be returned together with arsenic filter cake for smelting treatment so as to achieve the purpose of valuable metal recycling and solid waste reduction and harmless treatment.



Keywords Arsenic residue Neutralized residue Valuable metals Recovery





Gypsum

The waste acid generated during the copper smelting flue gas scrubbing and cleaning is characterized by high acidity. It not only contains As, F and other harmful elements, but also have Cu, Pb, Zn, Re and other metals. For the purpose to ensure the scrubbing effect of the dilute acid, it is necessary to discharge a part of dilute acid for treatment. For the widely applied treatment methods such as sulfide precipitation - gypsum for acid removal - lime and ferric salt neutralization, there exist deficiencies such as high treatment cost, more hazardous wastes difficult for treatment, valuable metals in waste acid not being recovered as resource, effluent quality hard to reach the requirement of reclamation etc. Along with the increasingly stringent requirement for environmental protection and the demand for the enterprise to establish a development model of cyclic economy, it is indispensable to research and develop new technologies of waste acid treatment with characteristics of resource comprehensive recovery. Y. Wen ⋅ Z. Bao (✉) ⋅ X. Wu Jinlong Copper Co. Ltd, Tongling 244021, Anhui, China e-mail: [email protected] © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_24

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This paper studies the feasibility for comprehensive recovery treatment of Cu, Pb, Zn, Re and other valuable metals based on the behaviors of waste acid generated during the copper smelting, probe into methods for reducing the impurities in flue gas in an effort to improve and optimize the dilute acid treatment process for reducing the output of arsenic filter cake, gypsum and neutralized residue, return the residue to the furnace for treatment and achieve the purpose of recovering the valuable metals in the waste acid and reduce the solid waste in a harmless way.

Waste Acid Treatment Process At present, a treatment process of sulfide precipitation - gypsum for acid removal lime and ferric salt neutralization is widely adopted for copper smelting waste acid [1]. Through precipitation of the waste acid, lead sulfate, silica and other insoluble substances. The supernatant is sent to sulfide precipitation procedure to generate sulfide deposits with heavy metals in waste acid and arsenic by using sodium sulfide, hydrogen sulfide and other vulcanizing agent. The sulfide deposits will be performed with solid and liquid separation by thickener, with arsenic residue and filtrate generated from press filter. The supernatant separated will enter the gypsum procedure. By adding limestone, it generates gypsum and calcium fluoride solid together with sulfuric acid and fluorinion. Through solid-liquid separation, gypsum is produced. The gypsum filtrate and field surface waste water containing copper and arsenic collected from sulfuric acid production will enter the lime and ferric salt neutralizing process via homogenizing pool. Add lime to adjust the waster water to the neutral condition, put ferrous sulfate in for oxidize the heavy metals and arsenic from low valence into high valence ions through air oxidation, feed lime for adjusting the pH value around 10 in order to produce metallic hydroxide, arsenate and calcium fluoride solids. Through precipitation by thickener, the underflow will be sent to the filter press for generating neutralized residue and the supernatant will be discharged up to standard [2].

Problems with Waste Acid Treatment Process and Its Reasons Problems with Waste Acid Treatment Process Conventional waste acid treatment method will generate a great number of arsenic residue, gypsum and neutralized residue. The residues contain Cu, Pb, Zn, Re and other valuable metals, which have not only increased the expenditure in solid waste

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Table 1 Composition of arsenic residue in Jinlong Copper Co Cu

S

As

Bi

Pb

Sb

Zn

Re

0.83

52.84

23.63

1.03

1.72

0.07

0.31

0.1

treatment but also resulted in waste of resources. To take Jinlong for example, the annual output of arsenic residue, neutralized residue and gypsum is 1,460 tons, 8,300 tons and 15,633 tons respectively. For composition of arsenic residue, please see Table 1:

Reasons for Large Quantity of Arsenic Residue The actual output of copper and arsenic filter cake is much more than the amount of theoretical calculation, which is mainly caused by four points: (1) SO2 contained in waste acid entering the vulcanization reacts with sodium sulfide during the vulcanizing to generate sulfur; (2) The waste acid contains suspended solids; (3) The sulfide deposits generated from valuable metal ions have entered the arsenic residue; (4) The content of arsenic waste in the waste acid is high [3].

Impact of SO2 in Waste Acid upon the Quantity of Arsenic Residue Based on the equilibrium constant ka = 0.78 between SO2 in the flue gas and and H2SO3 in the waste acid and the sulfite ionization equilibrium constant Ka1 = 1.54 × 10−2 and Ka2 = 1.02 × 10−7, the solubility of SO2 in waste acid containing 50 g/LH2SO4 is 12.53 g/L after calculation. SO2 solved in the dilute acid will enter into arsenic residue together with added sodium sulfide during the vulcanizing process, increasing the amount of arsenic residue. When the efficiency of absorption and desorption ranges from 0% to 100%, SO2 solved in 900m3 waste acid, Na2S consumed and sulfur increased in arsenic residue per day are as shown in the following Fig. 1: The above table shows that the efficiency of absorption and desorption for SO2 has a huge impact upon arsenic residue output. If the efficiency of absorption and desorption can increase by 376 ton dry residue per hour to the greatest extent, it is required to improve the efficiency of absorption and desorption to more than 99% so as to avoid additional residue and reduce the consumption of sodium sulfate.

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Fig. 1 Relation among efficiency of absorption and desorption, consumption of sodium sulfide and sulfur content in arsenic residue

Impact of Suspended Solids in Waste Acid upon the Quantity of Arsenic Residue and Loss of Valuable Metals The waste acid, after separation of suspended solids, enters the vulcanizing process. If the separation of suspended solids is not ideal in result, the suspended solids will go into arsenic residue after precipitation of vulcanizing process, resulting in increase of arsenic residue amount. The main component of suspended solids in waste acid is lead sulfate, so their access into arsenic residue will cause the loss of valuable metals such as lead, copper, bismuth, gold and silver. The composition of suspended solids is as shown in Table 2. Based on the current amount of waste acid, when the separation efficiency of suspended solids in the waste acid ranges from 0% to 100%, the amount change of the suspended solids entering the vulcanized arsenic residue is as shown in Fig. 2. Table 2 Composition of suspended solids in waste acid H2O

Cu

S

Bi

Pb

Au

Ag

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(%) 4.24

(%) 52.29

(g/t) 0.11

(g/t) 320.2

Fig. 2 Quantity of suspended solids accessing into vulcanized under different separating efficiency

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Table 3 Saturated concentration of As2O3 in gas of different temperature g/m3 500 °C

380 °C

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180 °C

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120 °C

100 °C

8840

3210

490

2.09

0.35

0.023

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Fig. 3 Quantity of metals accessing into vulcanized under different separating efficiency

The arsenic residue is expected to increase by 9 t/d to the maximum. Among others, the maximal amount of copper, bismuth and lead entrained into arsenic residue is 470.9 kg/d, 38.16 kg/d and 7.83 kg/d respectively. The quantity of vulcanized metals entrained for waste acid under different separation efficiency is as shown in Table 3. The conventional separation efficiency is only 80%, in this case, 0.56 t copper, 2.74 t bismuth and 33.9 t lead are entrained into arsenic residue each year. (Figure 3)

Impact of Arsenic Concentration in Waste Acid upon the Quantity of Arsenic Residue The arsenic in waste acid exists in the form of arsenious acid, with arsenic concentration determining the quantity of arsenic residue and consumption of sodium sulfide. The arsenic concentration is mainly affected by the temperature of smelting electrostatic precipitator. When the temperature of electrostatic precipitator is lower, arsenic trioxide will be condensed and the quantity of arsenic entering the sulfuric acid gas will be reduced. The saturated concentration of arsenic trioxide is as shown in Table 3. Table 3 shows that, when the flue gas temperature is reduced to 200 °C, 98% arsenic is collected by electrostatic precipitator. However, considering the corrosion to the equipment and flue by SO3 in flue gas, the flue gas temperature shall be more than 30 °C above SO3 dew-point temperature, therefore, the temperature of electrostatic precipitator has a great influence upon the arsenic residue.

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Influential Factor and Reason of Gypsum Quantity Most of arsenic and heavy metals in the waste acid generated during the copper smelting have been removed after vulcanizing reaction, however, there is still a small amount of arsenic and heavy metals entering the gypsum product during the gypsum making. Along with the stricter requirement of the state for environmental protection, the gypsum produced during the nonferrous metals smelting will be treated as solid waste. The sulfuric acid in the waste acid will react with limestone to form gypsum. The sulfuric acid in the waste acid comes from SO3, therefore, the generating rate of SO3 in flue gas has a direct impact upon the gypsum output. Based on the waste acid quantity of 900 m3/d in Jinlong, when the sulfuric acid concentration of the waste acid is between 1% ∼ 20%, the quantity of gypsum generated with water content of 10% is as shown in Table 4.

Influential Factors of Neutralized Residue Quantity Presently, most of smelters adopt lime-ferric salt process for treatment of waste water. This process usually generated a great number of neutralized residue with heavy metal elements. Among others, Zn and Cd entrained into smelting flue gas will mostly enter neutralized residue. Other small amount of As, Cu, Cr and other heavy metals will also enter the neutralized residue. In this case, the great number of neutralized residue has become a pollution source of heavy metals, making it difficult to dispose. The annual quantity of neutralized residue generated in Jinlong is 8,300 tons, with main components as shown in Table 5. Table 5 shows that the main components in neutralized residue are sulfur, calcium, iron, zinc, cadmium and other elements, among other valuable metals, the content of zinc and cadmium is 116.2 t and 11.6 t per year respectively. The main elements like sulfur and calcium in the neutralized residue are mainly sourced from sulfuric acid in the waste water and sulfate in added ferrous sulfate. Table 4 Quantity of gypsum generated with water content of 10% under different conditions of acid concentration t/d 1%

5%

10%

15%

20%

18

92

186

279

373

Table 5 Composition of neutralized residue % Cu

Pb

Zn

Cd

As

Fe

CaO

S

F

0.02

0.02

1.40

0.14

0.03

1.83

23.53

12.25

3.34

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Recovery of Valuable Metals in Waste Acid and Reduction and Harmless Treatment of Residue Recovery of Valuable Metals in Waste Acid Method for Recovery of Pb, Bi, Au, Ag and Other Valuable Metals PbSO4, Bi2O3, Au, Ag and Cu2S exist in the waste acid in a solid form, so, gravitational settling method is frequently used presently. Due to fine grain of PbSO4, the separating efficiency is not high and a part of Pb, Bi, Au, Ag are entrained into arsenic filter cake during the vulcanizing reaction of waste acid. Since the separating efficiency of suspended solids in waste acid by gravitational settling is low, it is imperative to develop a kind of separating equipment of membrane filtration with high separating efficiency and stable operation. Jinlong has developed a kind of cross-flow membrane filter, in which the feed liquid flows parallel to membrane surface. The shearing force arisen from feed liquid flowing across the membrane surface will carry the particles staying on the membrane surface away and keep the polluted layer in a thin level so as to solve the weakness of short sustainable operating duration and rapid flux decline, with the separating efficiency of suspended solids up to 99%. The separation by cross-flow filter allows to separate 99% solids containing Pb, Bi, Au and Ag from waste acid.

Recovery of Soluble Cu and Re After filtration, Cu and Re solved in the waste acid will step into vulcanizing process to form arsenic residue, which will not only increase the quantity of arsenic residue, but also cause the loss of copper, rhenium and other metals. In this case, it is possible to separate the sulfide and As by virtue of different equilibrium constants, settle the waste acid based on the feature of high binding power between sodium thiosulfate and copper ion and rhenium, with copper and rhenium producing rhenium-rich concentrate in a bid to achieve the goal of comprehensive recovery. Besides, the low precipitation of arsenic throughout the whole treatment process will exert no impact upon the subsequent treatment of waste acid [4]. Based on the current amount of waste acid of Jinlong, it is expected to recover 30 t copper and 1.5 t rhenium per year.

Recovery of Zn When the vulcanizing process under the conditions that dilute sulfuric acid is contained in the waste acid, it is not possible to generate zinc sulfide deposits. Through gypsum process in which pH value is controlled at 2.5, after most of

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Table 6 Weight and settling efficiency of zinc hydroxide under different pH conditions PH

6

6.5

7

7.5

8

8.5

9

9.5

Zinc Hydroxide kg/d Settling efficiency %

0.0

0.0

385.2

481.8

491.5

492.5

492.6

492.6

0.0

0.0

78.2

97.8

99.8

100

100

99.98

sulfuric acid is removed, it will step into neutralizing process. After neutralizing and adjusting pH value to 9, zinc hydroxide will be generated into neutralized residue. For recovering the zinc in the waste acid, it is possible to add Na2CO3 for adjusting the pH value and generating Zn(OH)2. To calculate based on the solubility product ksp = 7.1 × 10-18, the quantity and settling efficiency of zinc hydroxide can be obtained when pH value is 6-10, as shown in Table 6. Table 6 shows that when pH value is controlled at 8, 99.8% of zinc hydroxide will access into settling, with Zn(OH)2 as raw material recovered for sale.

Reduction and Harmless Treatment of Slag in Waste Acid Reduction and Harmless Treatment of Gypsum The sulfuric acid generating gypsum comes from SO3 in flue gas, therefore, the quantity of sulfuric acid in the waste acid can be reduced by reducing the quantity of SO3 entering the flue gas. As a part of residual oxygen is contained inside the flash furnace, it will produce a part of SO3. For reducing SO3, it is possible to make use of calcium oxide to react with SO3 for generating calcium sulfate. The decomposition temperature of calcium sulfate is 1,200 °C. Since the inlet temperature of the flash furnace boiler exceeds 1,200 °C, lower the temperature to 650 °C and make it enter the convection part, hence, it is easy to generate calcium sulfate by adding calcium oxide at the convection part of the boiler. If SO3 can be removed by 90%, 4.3 ton per day can be returned to the flash furnace for treatment, which will completely resolve the problem that a large quantity of gypsum cannot be disposed.

Reduction of Arsenic Residue After defining the four influential factors which results in actual output of arsenic residue more than theoretically calculated amount through the aforesaid analysis, four measures can be made to reduce the arsenic residue: (1) Improve the absorption and desorption efficiency of SO2 in waste acid; (2) Raise the separating efficiency of suspended solids in waste acid and reduce the solids entrained into

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vulcanization; (3) Separate the valuable metal ions copper and rhenium solved in the solution; (4) Reduce the temperature of electrostatic precipitator to make arsenic condense at the electrostatic precipitator and separate and reduce the content of arsenic entering the sulfuric acid flue gas [5].

Reduction and Harmless Treatment of Neutralized Residue Through vulcanization and gypsum process, most of heavy metals and H2SO4 in the waste acid have been treated. For small amount of heavy metals and H2SO4, the conventional lime-ferric salt neutralizing method is adopted by adding calcium hydroxide and ferrous sulfate to separate As and heavy metals in the form of arsenate and hydroxide. This method will generate a great number of neutralized residue containing heavy metals. For the solution after gypsum making, it is possible to adjust the pH value to 8 by adding sodium sulfate for generating sodium sulfate and solving in waste water, in this case, consideration can be given by adding sodium hydroxide and ferrous sulfate for oxidizing and generating ferrous arsenate deposits of heavy metals. As it has not formed into product of calcium sulfate, baaed on calculation of neutralized residue composition, CaSO4 • 2H2O makes up 72% of the neutralized residue quantity, in this case, the neutralized residue quantity can be reduced to 28% by changing the method of neutralizing agent, which can produce 2,320 tons per year, with 6.45 t neutralized residue per day on average. At present, there are certain smelters which have achieved good results in test of neutralized residue to be returned into the furnace for treatment [6], therefore, the neutralized residue containing arsenic and heavy metals can be returned to flash furnace for treatment.

Conclusion Through recovery of valuable metals from waste acid, 30.56 t copper, 33.9 t lead, 2.74 t bismuth, 1.5 t rhenium and 116 t zinc can be recovered annually. Through filtering of waste acid, 180 t arsenic residue can be reduced per year. By reducing the temperature of electrostatic precipitator, it is possible to reduce the arsenic residue. By filling calcium oxide into flash furnace boiler for reducing SO3, 1,400 t gypsum can be reduced per year. When the zinc is separated from the waste acid and the neutralizing agent is changed into sodium hydroxide, the quantity of neutralized residue can be reduced by 6,358 t per year. For the reduced gypsum and neutralized residue, due to low quantity on the whole, they can be returned into flash furnace for treatment, thereby achieving the harmless goal of residues.

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References 1. Yinglin P, Bin X (2016) Treatment of high arsenic content wastewater by two-step neutralization-iron salt precipitation. Ind Water Treat 36(6):64–68 2. Peng C, Shaopeng W, Panping X (2016) Discussion on problems of copper smelting waste acid treatment and their countermeasures. World Nonferrous Metals 8:76–77 3. Mian D, Zupeng L, Longwen C (2014) Practice of efficient vulcanization recovery of arsenic in waste acid treatment. Sulphuric Acid Ind 2:52–55 4. Yong-bin W, Jian-fen H, Wei L, Fu-ming L (2015) An experimental study of the enrichment of rhenium from copper smelting waste acid by the sodium thiosulfate precipitation method. Acta Petrologica et Mineral 1:110–116 5. Liang Y, Xu-peng L (2014) Research and application of quench arsenic collection technology for nonferrous metallurgy off-gas treatment. China Nonferrous Metall 3:41–44 6. Jufeng J (2016) Production practice of recycling neutralizing slag in Ausmelt smelting furnace. World Nonferrous Metals 2:30–31

Fundamental Process Equilibria of Base and Trace Elements in the DON Smelting of Various Nickel Concentrates Pekka Taskinen, Katri Avarmaa, Hannu Johto and Petri Latostenmaa

Abstract The converter-less nickel matte smelting technology (DON) adopted more than 20 years ago in Boliden Harjavalta smelter has been since that applied successfully to the processing of large number of nickel sulphide concentrates of various Ni-to-Cu ratios and MgO contents. The operational point of the technology is far from the conventional primary nickel smelting in the smelting-converting route. Therefore, a careful scouting of distribution equilibria of the base and trace elements in the smelting conditions of DON process has been conducted, in order to obtain quantitative information about the equilibria and thermodynamic properties of the nickel mattes at low iron concentrations, less than 10 wt% [Fe] in matte. The series of investigations has included novel experimental and analytical techniques for increasing the reliability and sensitivity of the phase equilibria as well as the element distribution observations carried out in typical high-grade nickel matte smelting conditions. Keywords Nickel matte Platinum group metals



Iron silicate



Magnesia



Precious metals

Introduction An access to high-MgO raw materials for nickel smelting directed Outokumpu in the late 1980s to development of a new primary smelting technology for nickel matte, based on the flash smelting concept. The fundamental novel idea was the expansion P. Taskinen (✉) ⋅ K. Avarmaa Department Chemical Engineering and Metallurgy, School of Chemical Engineering, Aalto University, P.O. Box 16100, 00076 Aalto, Finland e-mail: [email protected]fi H. Johto Outotec Research, Kuparitie 10, 28101 Pori, Finland P. Latostenmaa Boliden Harjavalta, Teollisuuskatu 1, 29200 Harjavalta, Finland © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_25

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of slag volume generated in the smelting step, by oxidation the feed mixture to much lower iron concentrations in the produced matte than the conventional practice [1]. This involved a new flow sheet, closing internal circulation of the slag and matte between the converting and slag cleaning thus improving the metal value recoveries, in particular that of cobalt, and lowering environmental impact of the nickel matte smelting. The elimination of converters and the entire converting step had several side effects to the industrial operation, including e.g. lower fugitive emissions and smaller CAPEX [2]. It also required modifications in the refinery flow sheet [3]. The current operational practices will be described elsewhere in this Conference [4]. Fundamentals of nickel smelting are much less scrutinised than those of copper smelting [2–5]. In 1995 only limited information existed, about the fundamentals of matte-slag equilibria, when the direct nickel matte smelting was taken into industrial use, and the data on properties of trace elements at low iron concentrations in the matte below 15 wt% Fe were non-existent. In high-iron mattes, the previous focus had been in nickel losses to slag and recoveries of selected trace elements, typically cobalt [6]. The knowledge in early ‘90 s upon nickel mattes in the converting was largely based on a review of Kellogg [7] and the prior experimental data. Font et al. [8] and Henao et al. [9] presented new experimental data on the slag-matte-gas equilibria in MgO crucibles and about selected trace elements. The scope was in conventional matte making and its conditions. Certain details on trace elements in nickel matte converting related to platinum group element distributions were studied [10]. For helping to understand the coupled phenomena in converting, process dynamics modelling has also been used [11, 12]. This presentation gives an overview on the recent studies of the slag-matte equilibria carried out with a novel experimental technique, allowing accurate observations about phase equilibria, their assays and the distributions of minority elements deporting them between matte to be recovered and slag where many elements will be discarded at low concentrations to various slag products.

Experimental The experimental technique was based on gas equilibration of small matte-slag samples on a solid substrate in flowing CO-CO2-SO2-Ar mixtures of controlled compositions. The experimental conditions were selected so that sulphur dioxide pressure in the furnace in all conditions was P(SO2) = 0.1 atm. This was a modification of a technique used earlier in many geochemical applications [13], and adopted by Jak et al. [14] for metallurgical slags and slag-metal equilibria. The experimental set up as well as the techniques used for confirming the state of equilibrium reached in the experiments of this work have been presented earlier in detail in the literature [15–17]. The experimental breakthrough in the analytical techniques for the trace elements was the use of laser ablation-inductively coupled plasma-mass spectrometry

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(LA-ICP-MS) directly from the polished sections, without separating the different phases of samples prior to the phase composition analyses. Combined with the electron microprobe X-ray analysis (EPMA) it allowed accurate chemical analyses in the whole range from several wt% to sub-ppm concentrations. The technique used also involved a statistical evaluation of the composition of each phase so that 8–10 points were measured on well-quenched domains of the sample and in addition to the average composition, its standard deviation will also be reported [15]. Those were the first metallurgical slag samples analysed with LA-ICP-MS technique and a lot of effort was put on the possible systematic errors arising from samples, from different composition domains than the geological specimens studied earlier [18]. An indication of the consistence of the two direct analytical methods is the good agreement for such elements present at concentrations above the detection limits of both the methods. An example is shown in Fig. 1 where cobalt concentrations of various iron silicate slags by LA-ICP-MS and EPMA are plotted as a function of iron concentration in the nickel matte at 1350–1450 °C. The agreement between two independent techniques is good, and the obtained standard deviation of the results is ± 0.01 wt% between the different techniques. The experimental series consisted of equilibration experiments at 1350–1450 °C for gas-matte-slag samples in fused quartz crucibles and with constant Ni-Cu ratios of 0, 2:1 and 4:1 (w/w). The studied sulphide mattes were synthetised in situ in the equilibration furnace from pure Cu2S, FeS and Ni3S2 powders. They initially contained 1 wt% of each trace element which were distributed between the slag, matte and gas phases during the high-temperature equilibration period. The main process between the gas, iron silicate slag and nickel-copper matte was the

Fig. 1 A comparison of EPMA and LA-ICP-MS data of cobalt concentration in iron silicate slags at nickel matte saturation containing about 0.5 wt% Co in different iron concentrations of the sulphide matte

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adjustment of iron distribution by two independent system variables, based on the overall reaction (1): ½FeSmatte + 1 ̸2O2 ðgÞ = ðFeOÞslag + 1 ̸ 2S2 ðgÞ.

ð1Þ

Thus, the prevailing sulphur and oxygen pressures as independent variables define the distribution of iron between the nickel matte and slag, as well as ‘the matte grade’ defined in this study as iron concentration of the sulphide matte. As a parallel reaction within the slag, iron oxides distribute between the oxidation stages as ðFeOÞ + 1 ̸4O2 ðgÞ = ðFeO1.5 Þ

ð2Þ

but the advancement of reaction (2) was not considered experimentally, as EPMA is sensitive to the elements only and no information about their oxidation states can be obtained.

Results The slag assay in matte-slag equilibrium at silica saturation was examined as a function of the matte grade and magnesia concentration. The EPMA results of iron concentrations at 1400 °C from 3 to 12 wt% are plotted on an isothermal constrained Gibbs triangle FeOx-MgO-SiO2 in P(O2) = 0.01 Pa in Fig. 2. The oxygen pressure range in the two mattes with [Ni]:[Cu] = 2 and 4 (w/w) had no major impact to the silica saturation boundary, as can be seen in the graph. It shows about 1 wt% higher silica solubility in contact with matte than the assessed copper- and nickel-free iron silicate slag system in the Mtox database [19] used in the calculations. The experimental liquidus composition data allow also estimation of the temperature dependency for the silica saturation boundary at fixed magnesia concentrations, Fig. 3. The used magnesia concentrations in the two experimental series with [Ni]:[Cu] = 4 and 2 (w/w) are not completely overlapping, in particular at the highest MgO concentration. The experimental technique used and the small differences sin the initial MgO compositions causes the recognisable scatter in the liquidus line projection shown in Fig. 3. The solubility of sulphur in the slag was studied as a function of temperature, iron concentration of matte and MgO concentration of the slag. MgO concentration of the slag has a clear effect on the sulphur solubility, as can be seen in Fig. 4. It is clearly also equally affected by the iron concentration of the matte in a fixed atmospheric point, which also reveals the activity of iron in the system. This indicates that the nickel-to-copper ratio has relatively small impact to the iron activity of matte in the current composition range.

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Fig. 2 Experimental slag composition of silica saturation at 1400°C superimposed on a calculated isothermal section of the system by MTDATA and Mtox database, vers. 8.2 [19]; CPX = clinopyroxene, HAL = halite (wüstite, MgO), OLI = olivine, OX_LIQ = slag, TRI = tridymite

The distribution coefficient of a component Me was defined in this study as ratio of its concentration in the nickel matte divided by that in the slag, i.e. Lm ̸s ðMeÞ ≡ ½wt − %Mematte ̸ ðwt − %MeÞslag

ð3Þ

The above referred fact that iron activity in nickel-copper-iron sulphide mattes at constant iron concentration is not strongly a function of the [Ni]:[Cu] -ratio is also seen in the behaviour of the distribution coefficient of iron at low iron concentrations, Fig. 5. Magnesia concentration of slag, affecting the iron activity of the slag in each oxygen pressure [17] and thus the matte composition in each equilibrium condition has been used as parameter in Fig. 5. The de-ironisation of nickel mattes proceeds thus in a similar way from mattes with a high and low copper concentration, over a wide range of [Ni]:[Cu] ratios, down to 2–3 wt% iron. The removal of iron from the nickel-copper mattes occurs at high iron concentrations essentially without major slagging of copper or nickel, and independently of the copper concentration of the matte, as scrutinised in industrial converter blows already by Browne [20].

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Fig. 3 Liquidus contour projection as a function of SiO2 concentration with MgO as parameter: silica solubility increases along with MgO-concentration; the calculated boundaries at MgO = 0 (–) and 9 wt% (―) are based on Mtox database (vers. 8.2) and MTDATA software

Fig. 4 Sulphur solubility in DON slags at iron concentrations of the matte from 3 to 12 wt%; the slags with a similar MgO-level at 1400°C have been plotted on the same graph and no impact of the [Ni]:[Cu] ratio of nickel matte can be seen (open symbols [Ni]:[Cu] = 2, closed [Ni]:[Cu] = 4)

The thermodynamic properties of nickel and copper in the matte vary when it is de-ironised in the smelting, and, as a consequence, when the Ni-Cu ratio of the feed mixture of the smelter fluctuates along with time. This causes changes in their

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Fig. 5 Distribution coefficient of iron between matte and slag as a function of iron concentration at 1400°C with MgO-concentration of the slag as parameter (open symbols [Ni]:[Cu] = 4, closed 2; ⃞ 8–9%, ⃟ 4 and ⃝ 0% MgO)

matte-to-slag distribution coefficients. That feature at 1400 °C is demonstrated in Fig. 6a and b for nickel and copper at (MgO) = 0, and in Fig. 7a, and b for nickel, at three magnesia concentrations of the slag. MgO additions to the slag favour the distribution of copper and nickel to the sulphide matte. The effect is small but clear in all studied concentrations of

Fig. 6 Distribution coefficient of nickel and copper between nickel matte and slag at 1400°C with different MgO concentrations in the slag (11–12 wt%, 6–7 wt% and 0 wt%: ●, ● and ○; [Ni]: [Cu] = 2)

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Fig. 7 The effect of [Ni]:[Cu]-ratio of matte on the distribution coefficient of nickel at (MgO) = 0 (a) and 8 wt% (b) at 1400°C, as a function of iron concentration of the matte (open symbols [Ni]: [Cu] = 4, closed 2)

magnesia. The distribution coefficient is affected slightly also by the used boundary condition of this study, when (Fe):(SiO2) ratio of the slag decreases with increasing MgO, as pointed out earlier, e.g., by Takeda [21] and Strengell et al. [16]. As suggested earlier (e.g., Teague et al. [6]) the reason of more favourable distributions between matte/metal and slag is the increase of the activity coefficient of less basic oxides, by formation of stronger MgO-SiO2 bonds in the slag when magnesia is added. The distribution coefficients of the trace elements cobalt and gold between the nickel-copper mattes as a function of its iron concentration and slag are shown in Figs. 9 and 10, respectively. The nickel-to-copper ratio of nickel sulphide matte in these studies was [Ni]:[Cu] = 4 (w/w). As can be concluded from Fig. 10, the solubility of palladium from the nickel mattes containing 1 wt% Pd in the iron silicate slag is very low (< 1 ppm). MgO clearly affects low concentrations of cobalt in iron silicate slags when upper iron concentration range of this study is concerned, see Fig. 8. Below ≈ 3 wt% iron in the matte, the impact of magnesia cannot be found any more. The MgO-free data obtained are in good agreement with Toscano and Utigard [22]. The PGM and PM distributions measured in copper (Cu-Fe) and nickel (Cu-Ni-Fe) matte-metal equilibria suggest that nickel mattes are more favourable collectors of precious and platinum group metals than copper mattes, as shown in Fig. 9. Nevertheless, the obtained distribution coefficients in both the cases are very high. The presence of basic oxides in the slag affects the distributions of the precious and platinum group metals very much. As Fig. 10 indicates, the impact of MgO on

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Fig. 8 The effect of MgO concentration in slag on the distribution of cobalt between nickel matte and iron silicate slag at 1400 °C in silica saturation; (wt% MgO) = 0, 4 and 8 and [Ni]:[Cu] = 4

Fig. 9 A comparison of matte-slag distribution coefficients for gold and platinum in copper and nickel matte-slag systems at 1350 °C and 1400 °C, respectively; [Ni]:[Cu] = 4 in the studied nickel sulphide matte

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Fig. 10 The effect of iron concentration in nickel sulphide matte and MgO concentration in the iron silicate slag on the distribution coefficient of gold at 1400 °C in silica saturation; [Ni]: [Cu] = 4, P(SO2) = 0.1 atm

the distribution coefficient of gold is about factor of 5 larger when its concentration in silica saturated iron silicate slag is increased from 0 to about 8 wt% (MgO). We did observe a similar trend in the distribution behaviours of palladium and platinum, as well [17]. Also here, increasing silica concentration of the iron silicate slag has a positive influence on the distributions [21].

Conclusions Due to the absence of data on nickel-copper-iron mattes in low iron concentrations, below 15 wt% [Fe] and in the operational window of DON process, an experimental program was carried out for measuring base metal and trace element distributions. The variables used were iron concentration of the matte, its nickel-to-copper ratio, magnesia concentration of the slag and temperature. Magnesia of the slag in these conditions, in a fixed atmosphere, has a clear impact to iron concentration of matte, as a feedback from iron activity of the slag [17]. The common substance in nickel sulphide concentrates’ gangue, MgO [24, 25], improves favourably recoveries of the base metals to the sulphide matte. A particularly large impact of MgO was found to be on those elements, which are weak oxide formers, as typically the precious and some platinum group metals.

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A comparison of the current observations at silica saturation with the computational phase diagram Fe-O-MgO-SiO2 based on Mtox database [19] indicates a good agreement. This suggests that the assessed data of Mtox database reproduces reliable phase property data for the industrial nickel matte smelting slags in DON smelting conditions in the flash smelting furnace (FSF) [26]. Font et al. [8] presented distribution data for selected trace elements in MgO crucibles at 1300 °C (i.e. in olivine saturation). They compared their observations as a function of P(SO2), with mattes of different [Ni]:[Cu] ratios and iron concentrations, including the limiting ‘binary’ matte systems Cu2S-FeS and Ni3S2-FeS. Their results indicate the effects of prevailing P(SO2) on the dissolution of the base metals into iron silicate slags. The impact of MgO on the distributions is not visible, due to olivine saturation where silica concentration of the slag still is a free variable, as indicated in Fig. 1, and was less accurately controlled in those experiments. As atmospheric SO2(g) links together the sulphur and oxygen pressures in the equilibrium systems. The high-SO2 environments thus represent higher oxygen pressures in a fixed matte composition, according to reaction ½Smatte + O2 ðgÞ = SO2 ðgÞ,

ð4Þ

where sulphur pressures in the slag-matte equilibria are controlled by the matte and its assay. Therefore, in the environments of flash smelting, when slag and matte are formed below the reaction shaft, on the FSF settler bath surface, from the oxidation products the settler reactions generate essentially pure sulphur dioxide gas and the local prevailing P(SO2) ≈ 1 atm [23]. This is also the universal boundary condition for the sulphide matte and slag forming reaction process in the flash smelting furnace settlers, independently of the oxygen enrichment. The results by Font et al. [8] imply that the matte-slag distribution coefficients for As and Sb increase when iron concentration of copper-nickel matte decrease, and the highest values for As, Bi and Sb were always obtained in copper-free mattes. This pattern, based on the present experimental observations, is more complicated if the effect of magnesia on the slag assay and its silica will be taken into account. There seems to be no data concerning the non-saturated magnesiabearing iron silicates. Acknowledgements The authors are indebted to Boliden Harjavalta Oy for its dedication and SIMP program by Tekes and Dimecc Oy for funding this extensive study.

References 1. Mäkinen T (1994) Method for producing high-grade nickel matte and metallized sulfide matte. US Patent 5 332 414. July 26, 1994 2. Mäkinen T, Taskinen P (1997) Physical chemistry of direct high-grade nickel matte smelting. In: 8th International Flash Smelting Symposium, Tucson (AZ)-Salt Lake City (UT). Outokumpu, Espoo, pp 251–295

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3. Mäkinen T, Taskinen P (2008) State of the art in nickel smelting: direct Outokumpu nickel technology. Trans IMM Sect C, 117(2):C86–94 4. Johto H, Latostenmaa P, Peuraniemi E, Osara K (2018) Review of operations at the Boliden Harjavalta nickel smelter. In: This conference 5. Diaz C, Landolt C, Vahed A, Warner AE, Taylor J (1988) A review of nickel pyrometallurgical operations. JOM 40(9):29–33 6. Teague K, Swinbourne D, Jahanshahi S (1998) Activity coefficient of cobalt oxide in non-ferrous smelting slags-a review. Aus. IMM Proceed. 303(1):1–6 7. Kellogg H (1987) Thermochemistry of nickel-matte converting. Can Metall Q 26(4):285–298 8. Font J, Hino M, Itagaki K (1999) Phase equilibrium and minor element distribution between iron-silicate based slag and nickel-copper-iron matte at 1573 K under high partial pressures of SO2. Mater Trans, JIM 40(1):20–26 9. Henao H, Hino M, Itagaki K (2002) Phase equilibrium between Ni-S melt and FeOx-SiO2 or FeOx-CaO based slag under controlled partial pressures. Mater Trans 43(9):2219–2227 10. Thyse E, Agdogan G, Eksteen J (2011) The effect of changes in iron-endpoint during Peirce-Smith converting on PGE-containing nickel converter matte mineralisation. Miner Eng 24(7):688–697 11. Kyllo AK, Richards GG, Marcuson SW (1992) A mathematical model of the nickel converter: Part II. Application and analysis of converter operation. Metall Trans B 23(5):573–582 12. Tan P, Zhang C (1997) Thermodynamic analysis of nickel smelting process. J Central South Univ 4(2):84–88 13. Kress V (1997) Thermochemistry of sulfide liquids I. The system O-S-Fe at 1 bar. Contr. Mineral Petrol 127(2):176–186 14. Jak E, Hayes P, Lee H (1995) Improved methodologies for the determination of high temperature phase equilibria. Metals Mater 1(1):1–8 15. Avarmaa K, O’Brien H, Johto H, Taskinen P (2015) Equilibrium distribution of precious metals between slag and copper matte at 1250–1350 °C. J Sustain Metall 1(3):216–228 16. Strengell D, Avarmaa K, Johto H, Taskinen P (2016) Distribution equilibria and slag chemistry of DON smelting. Can Metall Q 55(2):234–242 17. Piskunen P, Avarmaa K, O’Brien H, Klemettinen L, Johto H, Taskinen P (2018) Precious metals distributions in direct nickel matte smelting with low-copper mattes. Metall Mater Trans B 49(1):98–112 18. Jackson SE (2008) Calibration strategies for elemental analysis by LA-ICP-MS. In: Laser ablation–ICP–MS in the earth sciences: current practices and outstanding issues (P. Sylvester Edit.). Short Course Series #40; Mineral Assoc Canada, Quebec, pp 169–188 19. Gisby J, Taskinen P, Pihlasalo J, Li Z, Tyrer M, Pearce J, Avarmaa K, Björklund P, Davies H, Korpi M, Martin S, Pesonen L, Robinson J (2017) MTDATA and the prediction of phase equilibria in oxide systems: 30 years of industrial collaboration. Metall Mater Trans B 48 (1):91–98 20. Browne DH (1910) The behaviour of copper-matte and copper-nickel matte in the Bessemer converter. Trans AIME 42:285–305 21. Takeda Y (1997) Copper solubility in matte smelting slag. In: Proc molten slags, fluxes and salts ’97. Iron and steel society, Warrendale (PA), pp 329–339 22. Toscano P, Utigard T (2003) Nickel, copper, and cobalt slag losses during converting. Metall Mater Trans B 34(1):121–125 23. Taskinen P (2011) Direct-to-blister smelting of copper concentrates: the slag fluxing chemistry. Miner Proc Extr Metall 120(4):240–246 24. Taskinen P, Seppälä K, Laulumaa J, Poijärvi J (2001) Oxygen pressure in the Outokumpu flash smelting furnace—Part 2: the DON process. Trans IMM Sect C, 110 (2):C101–108 25. Mäkinen T, Taskinen P (2006) The state-of-the art in nickel smelting: direct Outokumpu nickel technology. In: Sohn Internat. Symposium, vol. 8. TMS, Warrendale (PA), pp 313–325 26. Mäkinen T, Fagerlund K, Anjala Y, Rosenback L (2005) Outokumpu technologies for efficient and environmentally sound nickel production. In: Nickel and Cobalt 2005, international symposium, CIM, Montreal, paper #8–5

Challenges and Opportunities of a Lead Smelting Process for Complex Feed Mixture Christoph Zschiesche, Mehmet Ayhan and Jürgen Antrekowitsch

Abstract Managing complex material streams within a lead smelting/reduction process has a long history at Aurubis. After modernization of the secondary lead smelter in 1991, Aurubis operates an electric furnace and a Peirce Smith converter to process complex secondary materials. This process links the copper and the lead metallurgy allowing Aurubis to properly combine primary and secondary process lines to optimize the processing of complex raw materials and smelter intermediates. However, combination of copper and lead metallurgy processing requires a continuous evaluation as raw material quality and composition varies. This paper describes Aurubis lead smelter processing and provides some fundamental considerations to optimize the lead processing, in particular with aspects related to speiss formation. Keywords Lead smelting metals Lead Copper





⋅ Complex feed mixture ⋅ ⋅ Sulfur

Recovery of valuable

Introduction Aurubis operations have been associated with the processing of primary and secondary materials for a long time. With the modernization of the copper smelter in 1972 by the adoption of the flash smelting technology, new challenges appeared, mostly related to impurity management. Then, with the acquisition of Hüttenwerke Kayser, in 1999, AG Aurubis extended his portfolio with a flowsheet for processing of secondary copper-lead-tin bearing materials (scrap, shredder material, residues, sludges, e-scrap). Following this step, primary copper business was further strengthen with the acquisition of Cumerio. Both acquisitions enabled Aurubis to become the world largest copper recycler and leading integrated copper producer. C. Zschiesche (✉) ⋅ M. Ayhan Aurubis AG (Research Development Innovation), Hamburg, Germany e-mail: [email protected] J. Antrekowitsch Montanuniversität Leoben (Non-ferrous Metallurgy Department), Leoben, Austria © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_26

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In addition to the primary and secondary copper assets, the ability to combine copper and lead metallurgy via a dedicated secondary lead smelter in Hamburg, allows Aurubis to have a more efficient processing of copper complex materials, producing other valuable metals such as Au, Ag, Bi, Sb, Sn, Se, Te, etc.). However, as complexity still growing, Aurubis has developed a new strategy conceptualized in its Vision 2025 to become a multi metal processor in order to ensure its future competitiveness. The future materialization of this vision imposes challenges to current available metallurgical flowsheets in order to be able to receive and process a broad range of minor element containing raw materials and operate with a high competitive metal recoveries. With the processing of complex feed material the complexity of the refining processes will also increase and require a high awareness for optimum use of the asset. To address the recent developments for the utilization of by-products (e.g. slags) and environmental legislations development of modified or new processes is essential. Flowsheet development also has to take into account the interrelation between of Cu, Pb and Zn metallurgies to optimize valuable metal recoveries and product quality. Aurubis is therefore committed to continue working in developing flowsheet solutions to combine in a more efficient way the capabilities of copper and lead metallurgy in order to optimize recovery of valuable metals. The current paper provides a general description of the current lead smelter flowsheet and discuss some key aspects required to optimize the speiss metallurgy.

Process Description Figure 1 shows the process flowsheet of the lead secondary copper smelter. The electric furnace as the center of the lead smelter links the Cu metallurgy with the Pb/ precious metals metallurgy of Aurubis. The electric furnace was commissioned in 1991; this was the replacement for the existing blast furnaces to improve the environmental aspect and reduce the off-gas volume due to heat generation by electricity instead of burning primary or secondary energy sources. That was the base to connect the process gas stream with the scrubber and cooling unit of the acid plant without influencing the quality of sulfuric acid production. The material preparation (sintering) was replaced by a crushing and pelletizing step. Since the commissioning, several modifications have been implemented to meet different requirements such as throughput increase, product quality improvement, improved emission management, life-cycle times and furnace availability [1, 2]. With its capability to process various secondary raw materials and intermediates the secondary smelting process is a key pillar of Aurubis recycling business on the market of Cu-, Pb- and precious metals-bearing materials [1]. The secondary smelter flowsheet allows recovery of precious metals by the combination of Cu and Pb smelting and refining processes. In this process Cu is recovered in a CuPb matte which is converted and recycled within the primary copper operations. The Pb

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Fig. 1 Aurubis lead smelting operations

phase captures Bi, Sb and Sn for recovery within the lead refining process. The speiss phase as As–Ni bearing material will be refined within the Primary Copper Smelter. The produced slag is utilized as construction material. Various intermediates of the lead refinery (dross), of the secondary smelting operations (converter slag, flue dust) and of the primary copper smelter/anode slime treatment process (flue dust, slag) will be treated within the electric furnace. Depending on the feed material the off-gas cleaning system is designed to manage fluctuating SO2 concentrations, Hg and halogen contents [2]. The new lead refinery was commissioned in 2014 and uses the well-known processes to extract the valuables into products like Te–salt, Sn–salt, PbSb–litharge and PbBi alloy. The precious metals will be extracted by adding Zn to produce in a rich bullion which is forwarded to the precious metals plant where also the precious metal stream from the primary production will be processed. The Cu-containing speiss has a high load of As, Sb, Sn and Ni which is processed in the primary smelters to use the existing impurity capacities in the Cu-anode. The produced slag ensures a low valuable metal contents, satisfying physical properties (viscosity, stability, etc) and requirements for utilization. The major intermediate is the produced flue dust which will be recirculated permanently. As can be noted from the presented process flowsheet, a key aspect of the lead secondary smelter is the processing of complex materials. This type of feed is the main source of different minor valuable metals, such as Ni, Sn, Sb, Bi, among others, important for Aurubis multi metal recovery strategy.

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Due to the nature of the lead smelting process, these elements will combine depending on the metallurgical operating conditions of the furnace generating intermetallic compounds, known as speiss. The formation of speiss has been a constant topic for metallurgist operating processes such as the secondary lead process in Hamburg. Some key aspects observed at Hamburg operations related to the speiss formation process are the following ones: • As is a key element in the formation of speiss, due to its ability to form intermetallic compounds with base metals such as Cu, Ni and Fe. So the main amounts of As will be within the speiss and blister copper phase. • Sb has a slightly different behavior as it will distribute between speiss and crude lead. Some Sb is also reporting to the blister copper phase. • Sn behaves in a similar way to Sb with a weaker tendency to distribute to the lead phase. Therefore special attention has to be given to understand the fundamental aspects associated to speiss formation.

Understanding Speiss Metallurgy Reviewing the publications during the last decades there is a tendency towards the investigation of fundamental aspects associated to speiss metallurgy. Since the fifty’s [3] the miscibility gap between matte and speiss has been discussed. Based on that the ternary or quasi-ternary diagrams e.g. As–X–S, As–X–X–S, As–X–Pb and As–X–Pb–S (X…Cu, Fe, Ni, Pb) were used to define different speiss types [3]. These results were used and linked with industrial applications (e.g. different campaigns within the blast furnace process, imperial smelting process) [4]. With a couple of investigations Gerlach, et.al. described the deportment of Ag between speiss and matte and issued the importance of As, Ni, Sb and Sn for the elemental behavior [5–8]. Following this various publications came from Japan [9–12] where the authors reviewed the recent results and focused on the matte/speiss and lead/ speiss equilibrium, especially on the Ag and Au deportment. Additional work was also done on the area of distribution equilibria between speiss and slag [11]. Investigations for minor element deportment between lead and iron speiss indicated that Fe more than As affects the minor element reporting [12]. Regarding the function of the speiss a series of experiments were performed to investigate the treatment of a copper matte under strong reducing conditions with additional metallic Fe to fix As in a FeAs alloy, collect the precious metals in the copper phase and capture e.g. Sb in the lead phase [13–15]. The transition and evaluation of this knowledge to an industrial scale is an aspect which is not covered that often. Nevertheless there are some publications [16–18] which present different approaches to extract speiss forming agents like As, Sb or

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Sn from their industrial processes. If these elements are in the feed mixture as sulphides or an excess of sulfur bearing agents will be provided then there is a possibility to extract them via the gas phase [16]. A major share of discussion is about the blast furnace process [17] or electric furnace process [18] and their characteristic speiss. In the scope is the minimization of precious metal losses and the fixation of As within the speiss. Due to the strong reducing conditions some metallic Fe can be present which will lead to a Fe-based alloy with high As amounts. As the literature indicates, the term speiss is used for a couple of materials differentiate in concentration and characteristic. In the most cases speiss is an undesired phase which collects also valuables like precious metals. Furthermore the control of analysis and quantity can be identified as challenge. Otherwise the speiss can be approached to extract impurities from a process based on the use of affinity differences of Cu, Fe, Pb, Ni and Co to As and S. Based on the compilation of available literature two publications [3, 4] are selected to discuss the relevance of speiss forming agents (As, Sb, Sn) in base metal processes (Cu, Fe, Ni, Pb). Both papers link fundamental work and help to understand the speiss formation with a couple of operational data. In the first one, Kleinheisterkamp (1948) reported the most relevant fundamental aspects about the miscibility of melts originating from CuPb process within a shaft furnace and delivered hints about how to use these fundamental information for practical work. The difference of a real speiss phase and an alloy is reasoned by the portion of total As + Sb + Sn to Co + Cu + Fe + Ni + Pb. The results showed that the affinity of Fe to As is stronger than it is the case for Cu to As which explain the formation of FeAs speiss (Fig. 2) in strong reducing conditions where metallic Fe is present. The role of Ni is an important one due to its larger affinity to As than for Cu. Following this it seems to be possible to control the Ni deportment within a CuPb process if a certain level of As is in the feed. The solubility of metallic Cu and sulfidic Cu in Ni-arsenides complicates the control of this. The formation of Fig. 2 Ternary phase diagram As–Fe–S (T = 1180– 1320 °C) [3]

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Ni-arsenides is strongly affected by the As and S content in the feed mixture. The S content in the feed determines the Cu extraction and ensure the formation of matte. The more efficient this take place the less metallic Cu is present to form intermetallic compounds (e.g. Cu antimonides) which will increase the total wt% Cu in the speiss. [3]. In the second one, Fontainas (1978) discussed how to deal with the treatment of complex materials containing lead, copper, nickel, zinc, iron and sulphur. An important part of this is the discussion about various forms of speiss in the system slag, matte, speiss and crude lead. The author agreed on the description of Kleinheisterkamp and also reported about the broad manner of composition (Fig. 3) instead of a particular point for speiss composition. Furthermore the paper took into account that a high enriched metallic copper phase will appear if the activity of Cu is nearly unity but in reality it will be below 0.3. In combination with present impurities like As, Sb or Sn the Cu will form intermetallic compounds which are known as Cu-rich speiss. The formation of an iron-rich alloy (dominated by FeAs or Fe2As) mainly depends on the amount and form of compound of Fe in the feed mixture, the amount of the other minors and the chosen process atmosphere which is characterized also by final Pb content in the slag. An exemplary reduction process with metallic Fe as reduction agent and decisive amount of minor elements in the feed will produce a low grade CuPb matte with high amounts of Fe and a speiss which can contain up to 60 wt% Fe and 20 wt% As. From a stoichiometric point of view this ratio gives an indication for the dominated phase in the speiss (Fe2As). The referred blast furnace process (Fig. 4) will end with 1.5 wt% Pb which means a Fe activity of about 0.1. Under these circumstances the formation of an FeAs main alloy seems to be negligible but this turns when the Fe activity increases due to deeper reduction at least for an activity of 0.5 (Imperial Smelting process) which will promote the formation of a ferrous speiss. Beyond As also Sb and Sn

Fig. 3 Cu–Ni–As-based alloys and speiss [4]

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will form intermetallic compounds and become a part of the speiss. In Fig. 6 three different operations with speiss production are highlighted. Starting with a low Cu activity of 0.15 in the processing of Pb-bearing materials these speiss are different from the ones produced during the processing of Pb-Cu complex mixture. Then the activity of Cu is increased and this will lead to an excess of copper in the speiss composition. The current speiss composition of Aurubis lead smelting process in the electric arc furnace is in-line with the so called area Pb–Cu charges. If the Cu activity further increases Cu rich alloys will be formed in Cu-charges and this can continue until the right corner is reached where a black copper composition will be reached. With the help of Fig. 4 also the role of Ni will be explained by applying a ternary Cu–Fe–As diagram in which Ni is lumped together with Cu, Sb, Sn and As and with Co, Fe and Zn. Ni is to consider as a base metal like Cu which will form intermetallic compounds e.g. Ni5As2, NiAs, NiSb. The Aurubis lead smelting process is close to the iso-activity curve for a typical blast furnace process (with 1.5 wt% Pb in slag). At this point the formation of Fe-rich alloys is not expected due to the low activity of Fe (0.1) [4]. Figure 5 issues the deportment of As, Sb and Sn between crude lead and speiss phase. To keep it simple the diagram relates on a fixed temperature (T = 600 °C) where the solidifcation of speiss is assumed as completed and the lead is still in liquid state. Based on ternary diagrams of Pb–Cu–As, Pb–Cu–Sb and Pb–Cu–Sn Fig. 4 compile all information for As, Sb and Sn concentration in lead versus the concentration of these elements in the Cu-alloy (speiss). This means for a Cu–As alloy with 30 at.% As the concentration in the lead bullion will be about 0.45 wt% As. For As it can be stated that the As deportment is not that strong related to the As content in the Cu-alloy like it is the case for Sn and Sb. Equal amounts of Sn in

Fig. 4 Cu–Fe–As-based alloys and speiss [4]

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Fig. 5 As, Sb and Sn content of lead bullion as a function of composition of Cu alloy/speiss in equilibrium at 600 °C speiss [4]

Cu-alloy (30 at.% Sn) will lead to 6 wt% Sn in lead (in equilibrium) [4]. Sn and Sb show a similar behavior in terms of element deportment. The following conclusions can be derived from this literature review: • Formation of intermetallic compounds will be governed by a feed mixture which contain base metals (Co, Cu, Fe, Ni) and minor metals (As, Sb, Sn). • Depending on reduction potential, used capability of Cu extraction and ratio of base metals to minor metals will be determining factors which speiss will be formed (quantity and quality). • Fundamental work deal with ternary and quasi-ternary systems to investigate the miscibility gaps between matte and speiss and the impurity behavior between speiss and lead. • As minor element the authors mean also Ag and Au whose deportment between speiss and matte and speiss and lead is investigated in a couple of papers. The formation of a Fe/As-bearing speiss will reduce the solubility for Ag which then can be collected in the matte or metallic phase. • The thermodynamic prediction for Ni is not that easy due to a lack of data in all relevant databases. Further fundamental experiments are required to address the importance of Ni for speiss formation. • To take Ni into account the phase diagrams (Figs. 3 and 4) can be used to calculate the expected speiss composition for a given copper activity or describe an operational window when the speiss tends to become Fe rich (for higher Fe activity and corresponding low lead content in end slag).

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Fig. 6 Dependency of sulfur source addition on the matte grade (top); Dependency of sulfur source on the Cu content in the metal phase (bottom) at T=1150 °C [20]

• Major share of available data is originating from equilibrium tests in lab-scale or bench scale tests. This means the use of synthesized material and the simplification of conditions. Industrial data of various speiss types is quite rare and if present then it will come from a couple of years ago. • Figure 5 shows a compilation of empiric data which allows the calculation of the As, Sb and Sn content in the crude lead following a complete solidification of speiss. In addition, the following implifications for industrial lead smelting process can be derived from the above analysis: • In a couple of literature sources (e.g. Kleinheisterkamp [3]) it is indicated that it is possible to properly determine potential formation of NixAsy speiss on its different stoichiometry. • The Fe content of the speiss depends on the presence of metallic Fe which can be minimized by controlling the reduction potential of the process.

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• A proper Cu extraction by decreasing the Cu activity will decrease the Cu content in the speiss because of less intermetallic Cu compounds with As, Sb and Sn. • Although the economic improvement is a main criteria for lead smelting processes, the operation with campaigns would be a serious solution to react on varying feed composition and provide a suitable product quality of lead, matte and speiss.

Approach for a Continuous Improvement of the Secondary Lead Smelting Process The main purpose of Aurubis Cu/Pb process is the recovery of precious metals in the crude lead phase and the extraction of copper via matte formation. On the other hand, the extraction of minor elements (e.g. As and Ni or Sb and Sn) has to be adjusted to achieve a high efficient process. All in all this will support the recovery of valuables and ensure the use of crude lead which can handle Sb and Sn but also the use of the other intermediates (matte for Cu and the speiss). In order to optimize this aspect, special attention has to be given to: • Calculation of feed material and scheduling the operations to address the fluctuating feed composition • Monitoring of product analysis to evaluate key factors for the feed (e.g. As, Sb, Sn and Ni content) • Proper sulfur adjustment to capture the copper and keep the matte grade above 30 wt% • Avoid a limited solubility for Sb and Sn in lead due to an insufficient reduction potential or less metallic Pb amounts

Possibilities to Adjust the Right Metallurgy for Recovery of Cu Within the Secondary Lead Smelting The Cu/Pb separation depends on the sulfur adjustment in terms of quantity and procedure. Sulfur can be added from different sources (elemental sulfur, pyrite or a lead concentrate), however each one of these sources will have a different efficiency, as discussed later in this section. • Elemental sulfur addition: Because of the low melting point (115.2 °C) and the low boiling point (444.6 °C) elemental sulfur is not favored due to its low efficiency [19]. • Pyrite/Lead concentrate: This has been industrially proven [3].

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Beyond these both additives additional sulfur sources will enter the secondary lead smelting process and support the formation of CuPb matte. The following reactions scheme introduce the principles of the mentioned process. Sulfates will be reduced directly by coke or indirect by carbon monoxide (Eqs. 1 and 2) and will provide PbS which then allows the sulfidation of Cu to form Cu2S (Eqs. 3 and 4). The advantage of using lead concentrate is the supply of additional metallic lead which is explained in Eq. 5. This additional lead allows the capturing of Bi, Sb and Sn. The major share of metallic lead will be originating from a lead oxide slag which is reduced (Eqs. 6 and 7). Addition of pyrite will enable the reaction with Cu2O to Cu2S which is mentioned in Eq. 8. This will also effect the matte grade due to additional amounts of FeS in the matte (Eq. 9). 2½C  + fO2 g = 2fCOg

ð1Þ

½PbSO4  + 2½C  = ½PbS + 2fCO2 g

ð2Þ

½PbSO4  + 4fCOg = ½PbS + 4fCO2 g

ð3Þ

½PbS + ðCu2 OÞ = ½Cu2 S + ðPbOÞ

ð4Þ

3½PbS + 2ðCu2 OÞ = 2½Cu2 S + 3½Pb + fSO2 g

ð5Þ

2ðPbOÞ + ½C  = 2½Pb + fCO2 g

ð6Þ

ðPbOÞ + fCOg = ½Pb + fCO2 g

ð7Þ

3½FeS2  + 2ðCu2 OÞ = 3½FeS + 2½Cu2 S + fSO2 g

ð8Þ

½FeS + ðCu2 OÞ = ðFeOÞ + ½Cu2 S

ð9Þ

Taken the low efficiency of elemental sulphur into account the use of pyrite is an alternative which provide an improved efficiency of S use due to the smaller reaction surface. For lead concentrate a significant higher amount of material is required due to the low share of S in lead concentrate. For the reaction with one tons of Cu the following amounts are required: • 800 kg elemental sulphur (assumed efficiency of 25%); disadvantage → a lot of SO2 emissions • 1850 kg lead concentrate • 930 kg pyrite (following the reaction in the scheme—Eq. 8) The effect of using pyrite or lead concentrate will be discussed in the following thermodynamic calculations which were based on a given example by applying a feed mix which is dominated by a high amount of Pb (20–40 wt%). Other base metals like Cu and Fe are also present. Because of the lack of data Ni was not considered. The following fixed parameters (process temperature 1150 °C, closed

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system without any gas exchange, equilibrium between slag/matte/metal) were set for the calculations. As scenarios the following cases were calculated: • I—Base case applies the feeding of lead concentrate and pyrite in a ratio of (11:1) by adjusting 2 wt% S in the feed • II and III—Increase of lead concentrate amount by 25% steps whereas the second steps results in 2.3 wt% S in the feed • IV—Decrease of lead concentrate amount by 50% in one step results in 1.7 wt% S in the feed • V Replace the lead concentrate in base case by pyrite to adjust the same 2 wt% S in the feed • VI Apply case V and reduce the amount of pyrite which means 1.2 wt% S in the feed The first plot in Fig. 6 shows the dependency of matte grade on the oxygen potential and the second one discusses the corresponding Cu content in the metal phase. The “worst” case (1.2 wt% S) means an insufficient sulfidation which will lead to an increase of Cu content in the metal phase. Cu which is not bound on sulfur is able to form intermetallic compounds by capturing As, Sb or Sn. On the other site an exceed of sulfur in the feed (pyrite and lead concentrate like in base case +50% more lead concentrate) will lead to a poor Cu content in CuPb matte with contents of 25 wt% Cu. But at the same time the Cu which is present in the metal will also shrink and concludes to a less amount of intermetallic copper compounds, better known as speiss. The dependency of the matte grade on the present oxygen potential can be specified for log pO2 from −10 to −11 with 27.5–30 wt% Cu (base case, T = 1150 °C). If lead concentrate is substituted by pyrite which will provide the same amount of introduced sulfur the matte grade will be effected and changes from 29.6 to 28.8 wt% Cu. In contrast to the effect on the matte grade for the metal phase an increase of Cu concentration is associated with increasing reduction potential. An exceed of sulfur will decrease the Cu content in the metal whereas a lack of sulfur increases the Cu content. The results allow the following interpretation: • In principle both additives will address the task to deliver sulphur to the system to form a matte for Cu extraction. • The expected matte grade is influenced by the reduction potential. For a constant log pO2 the sulfur adjustment is the main paramter which is changing the matte grade. • A well-calculated sulfur adjustment will determine the operational window for the matte and will have also an effect on the metal phase composition. • Contradictory to the matte grade change the Cu content in the metal will increase with higher reduction potential. • Which sulfur adjustment will be the right one depends on a couple of parameters and has to be defined from one to another operation. The following processing step for the matte has to be considered as well.

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Conclusion In the present paper Aurubis lead smelting process and its main function was presented. For the upcoming challenges operating a flowsheet which is able to link the Cu metallurgy with the complex metallurgy will be a key aspect. Therefore the most relevant criteria is the achievement of high metal recovery rates for a couple of valuables and the precious metals. The feed of the electric furnace will become more complex and this complexity will be somehow transmitted to the products (CuPb matte, crude lead bullion, speiss and a CaO–FeO–SiO2 based) and the following refining steps. The metallurgy has to be managed actively to react on the feed composition and obtain the right operational windows to provide a proper product composition for their foreseen treatment within Aurubis. The formation of intermetallic compounds in a separate phase, better known as speiss, is still not well understood. Fundamental work is still required to support proper metallurgical operations. Development of thermodynamic data base will support this goal. Secondary copper lead metallurgy requires proper understanding of required Cu, Pb and S ratios to optimize the metallurgical process and recovery of valuable metals. The extraction of Cu will be controlled by the addition of a suitable sulfur source which can be lead sulphate rich dusts, lead sulphate based slimes, lead concentrate or pyrite. The last two additives were discussed in the paper. The transition of the mentioned knowledge will support the aim to improve the impurity management of Aurubis and will provide a great future for the lead smelting activities within the electric furnace. With the right know-how the process will be fit for purpose and deliver it’s contribution for increase of metal recovery rates and more intelligent impurity control. Acknowledgements The authors gratefully acknowledge the management of Aurubis for the possibility to work on this particular topic and the permission to publish this. They would like also to thanks the colleagues who contributed to the comprehensive discussions. Special thanks to Dr. Mehmet Ayhan who fed his expertise to the work and raised a lot of challenging questions. The detailed review of Dr. Gerardo R. F. Alvear Flores helped to significantly improve the original version of this manuscript.

References 1. Bauer I, Hoffmeister F (2013) Increased productivity of the secondary electric furnace at Aurubis hamburg. In: EMC 2013, vol 2, pp 509–518 2. Bauer I, Kadereit H (2011) New off-gas treatment facility in the secondary smelter of Aurubis AG in hamburg. In: EMC 2011 vol 1, pp 199–208 3. Kleinheisterkamp H (1948) Einige grundlegende Gleichgewichte bei der Bildung metallurgischer Speisen, Zeitschrift für Erzbergbau und Metallhüttenwesen, Heft 3, S. 65–72, Juni 1948

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4. Fontainas L, Coussement M, Maes R (1978) Some metallurgical principles in the smelting of complex materials. Complex Metall 1978(1978):13–23 5. Gerlach J (1966) Über Untersuchungen im System Cu-As-S und die Verteilung von Silber auf Stein und Speise. Erzmetall, Heft 9:458–463 6. Diettrich H, Gerlach J (1967) Über Untersuchungen des Systems Cu-Ni-As-S und die Silberverteilung auf Stein und Speise. Erzmetall Heft 7:308–314 7. Gerlach J, Hennig U, Siewert H (1969) Die Silberverteilung auf Stein und Speise, Erzmetall, Band 22. Heft 12:586–591 8. Gerlach J, Hennig U, Trettin K (1969) Die Silberverteilung auf Stein und Speise, Erzmetall, Band 22. Heft 3:119–122 9. Espeleta AK (1991) Thermodynamic properties and phase equilibria of lead smelting speiss systems. Doctoral Thesis of Tohoku University, Institute of Mineral Dressing and Metallurgy 10. Hino M (1998) Speiss formation and phase equilibria. Shigen-to-sozai, Japan, 114(4):215– 223 11. Mendoza G, Hino M, Itagakai K (2002) Distribution equilibria between Cu-Fe-As ternary speiss and slag phases. Shigen-to-sozai, Japan 118:197–201 12. Hino M (1995) Distribution of Minor Elements between Molten Lead and Iron Speiss, Zinc & Lead ‘95. In: Proceedings of the international symposium on the extraction and applications of Zinc and Lead, Sendai, pp 737–746 13. Itagaki K, Voisin L, Mendoza D (2006) Phase Relations Copper Smelting Cu-Fe-S-X and Cu-Fe-S-C-X (X=As or Sb) Systems and Distribution of Precious Metals Relating to Reduction Smelting Copper. In: Sohn International symposium 2006, vol 1, pp 289–300 14. Damm G, Voisin L (2013) Phase relations and minor element distribution in Cu-Fe-Pb-As System saturated with carbon at 1473 K. Copper 2013:1153–1164 15. Voisin L (2013) Phase relations and minor element distribution in Cu-Fe-Pb-Sb System. Proc Copper 2013:1165–1179 16. Dosmukhamedov N (2016) Efficient removal of arsenic and antimony during blast furnace smelting of lead-containing materials. Miner Metals Mater Soc J Metals, Published online (2016) 17. Chaidez-Felix J (2014) Effect of copper, sulfur, arsenic and antimony on silver distribution in phases of lead blast furnace. Trans Nonferrous Metals Soc China 24:1202–1209 18. Fontainas LM (1980) A two-step process from smelting complex Pb-Cu-Zn-materials. Lead-Zinc 1980:375–393 19. Wikipedia, Webpage (2018). https://en.wikipedia.org/wiki/Sulfur 20. Jak E et al C153 database FactSage, University of Queensland

Application of MPE Model to Nickel Smelting Chunlin Chen

Abstract Nickel-sulfide minerals are normally associated with copper and iron sulfides and often contain a minor amount of valuable metals such as cobalt and detrimental impurities such as arsenic. Metal losses in slag vary from process to process during the pyrometallurgical production of nickel and, depend on the feed composition, slag chemistry and operating conditions. Maximiz the metal recovery is one of major considerations to optimize operating condition for nickel smelting/ converting processes. At the same time, the deportment of minor elements between various phases during nickel smelting is of great importance by smelter operators. The Multi-Phase Equilibrium (MPE) is a thermodynamic package developed by CSIRO for simulating reactions between phases in multi-component and multi-phase systems [1]. Over the years the capability of the MPE model has been extended to cover the behavior of a large number of elements in high temperature systems. The sulfide smelting module of the MPE, which covers the minor elements such as As, Bi, Sb, Pb, Se, Te, Sn, Co and Zn, is capable of modeling the deportment of major and minor elements between various phases during nickel smelting. In this article the application of the MPE model in modelling the nickel smelting process is presented. The modelling results on the majors and minors are compared with the plant data. The deportment behaviour of arsenic in nickel smelting was analysed and the impact of slag chemistry on slag properties was modelled. These results can assist process metallurgists in exploring the optimum fluxing strategy for smooth operation and the potential practices for improving arsenic removal. Keywords Nickel smelting



Thermodynamic modelling



Minor elements

C. Chen (✉) CSIRO Mineral Resources, Private Bag 10, Clayton South, VIC 3169, Australia e-mail: [email protected] © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_27

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Introduction Metal losses in slag vary from furnace to furnace during the pyrometallurgical production of nickel, depending on the feed composition, matte/slag chemistry and operating conditions. Minimizing metal losses is one of the major factors in considering the optimal operating condition for nickel smelting/converting processes. At the same time, the deportment of minor elements between various phases during nickel smelting/ converting is of great importance to smelter operators. This is driven by the requirement of control of product quality and minimizing the environment impact. To assist process metallurgists in assessing the impact of smelting nickel concentrates with various Ni/Cu ratio, MgO content and high levels of minor elements, an advanced thermodynamic model of the base metals smelting process which is developed from the process fundamentals and allows the deportment of major/ minor elements in the smelter under various operating conditions is required. Since the early 1990’s, CSIRO has been involved in development and application of thermodynamic models of melts and solid solutions. Over the years the capability of the CSIRO’s Multi-Phase Equilibrium (MPE) model has been extended to cover the behavior of a large number of elements in high temperature system [1]. The copper smelting module of the MPE, which covers the minor elements such as As, Bi, Sb, Pb, Se, Te, Co, Sn and Zn, as well as some naturally occurring radio-nuclide elements, is capable of modeling the deportment of major/minor elements in the smelter. Our previous published work has provided some details of the assessment carried out as well as application of the model to predict deportment of some minor elements during the smelting of sulfide concentrates [2, 3]. In the present paper, a series of plots covering the validation of the model and databases relevant to nickel smelting against published experimental data are presented. Application of the model in predicting the metal losses in slag, slag properties, as well as the arsenic deportment between matte, slag and gas phases of nickel flash smelting have been demonstrated. These results can aid process metallurgists in understanding the slag properties and the arsenic deportment behaviour during nickel smelting and exploring the potential practices for improving arsenic removal and process performance.

Thermodynamic Models and Database Ni-Cu Alloy Phase Unlike copper smelting, nickel sulfide converting does not produce nickel metal and the process stops at the slag-blowing stage to produce a high grade nickel matte which is used for electrorefining. The Ni-Cu alloy system is important for applications in nickel smelting, since the Ni-Cu system was often used as a reference for activity measurement of Ni in slag and matte systems.

Application of MPE Model to Nickel Smelting

(a) 1600

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Liquid

Activity, Ni

Temperature/ oC

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Schuermann & Schulz [5] Feest & Doherty [6] Predel & Mohs [7]

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0.8

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0.0 0.0

0.2

0.4

0.6

0.8

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X(Ni)

Fig. 1 Comparison of a the calculated Ni-Cu phase diagram and b nickel activity with experimental measurements. Lines are calculated by model. Symbols are experimental data [5–8]

The Redlich-Kister-Muggianu polynomial [4] is used in the MPE package to describe the excess Gibbs energy of the Ni-Cu alloy phase. Ni-Cu is a very well-known isomorphous system. The liquid phase is miscible in all compositions. The face centred cubic (fcc) (Cu, Ni) solid solution is also miscible down to the critical (Tc) temperature where it shows immiscibility over a wide range. Recent studies on the phase equilibria of the Ni-Cu system were done by [5–7] using XRD and microstructural analysis. Ni activities in the liquid Ni-Cu system at 1400 °C were measured by using EMF method by Kulkarni and Johnson [8]. As shown in Fig. 1, the Ni activities shows positive deviation. Figure 1 shows that the experimental measured Ni-Cu phase diagram and activity data [5–8] can be represented very well by MPE.

Slag Phase The molten slag phase is described by the Cell model [9] in MPE. The slag database includes the following species: FeO, Fe2O3, SiO2, CaO, MgO, Al2O3, MnO, Cu2O, CuO, NiO, PbO, ZnO, P2O5, TiO2, Ti2O3, CrO, Cr2O3, CoO, B2O3, SnO, As2O3, Bi2O3, Sb2O3 as well as S, Se, Te, Cl and F as anions. Takeda et al. [10] measured the distribution of Ni between the calcium ferrite slag and Ni-Cu alloy at 1250 °C and oxygen partial pressure of 10−9 to 10−5 atm. The Ni distribution between magnesia saturated iron silicate slag and Ni-Cu alloy at oxygen partial pressure range of 10−10 to 10−6 atm and temperature of 1400 and 1500 °C were reported by Pagador et al. [11]. The distribution data were plotted against the oxygen partial pressure in Fig. 2a. The linear relationship between LogD (s/m) (Ni) and the oxygen potential shows a slope of 1/2. This suggests that the

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C. Chen

(a)

(b) 4.0 T=1300 oC PSO2=0.1 atm

o

Iron silicate slag, T=1400 C Calcium ferrite slag, T=1250 oC

3.5

Cu/Ni=1

3.0 2.5

Ni in slag, wt%

slag/cu

1

D

Ni

0.1

0.01

2.0 1.5 1.0 0.5

1E-3

1E-10 1E-9

1E-8

1E-7

1E-6

0.0 10

1E-5

PO2, atm

(c)

20

30

40

50

60

70

80

(Cu+Ni), wt%

(d)

4.0 T=1300 oC PSO2=0.1 atm

3.5

Cu/Ni=1

0.1

2.5

DAsSlag/Cu

S in slag, wt%

3.0

2.0 1.5

Takeda et al., 1250 oC Calcium ferrite slag Kashima et al., 1300 oC Iron silicate slag calculated

0.01

1E-3

1.0 1E-4

0.5 0.0 10

20

30

40

50

(Cu+Ni), wt%

60

70

80

1E-5 1E-12 1E-11 1E-10

1E-9

1E-8

1E-7

1E-6

PO2, atm

Fig. 2 Comparison of the calculated a Ni distribution between Ni-Cu alloy and both iron silicate slag and calcium ferrite slag; b Ni solubility in iron silicate slag in equilibrium with Cu-Ni matte; c S solubility in slag in equilibrium with Cu-Ni matte and d arsenic distribution between slag and liquid copper with experimental measurements. Lines are calculated by model. Symbols are experimental data [10–12]

predominant species of Ni in slag is NiO and that the experimental data can be represented well by the model. Font et al. [12] equilibrated iron silicate slag with the Cu2S-Ni2S3-FeS matte with Cu/Ni of 1 under various PSO2 at 1300 °C. It was reported that the nickel content in the slag increases slightly in the region of matte grade up to about 50% and remarkably above this grade with increasing matte grade due to the sharp increase of oxygen partial pressure, as shown in Fig. 2b. Figure 2c shows that the sulphur content in slag decreases rapidly with increasing matte grade. It’s demonstrated in Figs. 2b, c that the experimental measured Ni and S content in the slag equilibrated with Cu-Ni matte can be represented by the model.

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The arsenic distribution between liquid copper and both iron silicate slag and calcium ferrite slag under various oxygen partial pressures at 1250 °C were reported by Takeda et al. [10]. The comparison of the model calculated As distribution ratio with the literature data [10] are shown in Fig. 2d.

Matte Phase The Quasi chemical matte model developed by Kongoli et al. [13] for the Cu-Fe-Ni-Co-S matte was adopted in the MPE package. This model was extended to cover the minor elements in the matte by introducing the cation–ME (minor element) interactions and the ME–sulfur interactions. The current matte database covers the following elements: Cu, Fe, Ni, Co, Cr, S, O, As, Sb, Bi, Pb, Zn, Se, Te and Zn [2]. Font et al. [12] determined the matte and slag chemistries from samples that were equilibrated under controlled SO2 and S2 partial pressures. In their experiments, the SO2 and S2 partial pressure were controlled by passing Ar-SO2 gas mixture through a reservoir of liquid sulphur, which was maintained at a constant temperature. The results by Font et al. [12] were used to validate the MPE. Figure 3a shows that the calculated S2 partial pressures of the copper matte equilibrated with iron silicate slag is slightly higher than the experimental value. Figure 3b shows that the calculated As distribution ratio are higher than the measured value [12], but the trend of decrease of As distribution ratio with increasing matte grade can be represented by the model.

(a)

1

(b)

0.1

Iron silicate slag o T=1300 C; NNi/NCu=1.0

1

PSO =0.1 atm 2 PSO =0.5 atm 2 PSO =1 atm 2

0.01 PS2

1E-5

P

SO

2

=0.5

atm

s/m

1E-4

o

T=1300 C PSO2=0.1 atm

P

LAs

Pressure, atm

1E-3

Cu/Ni=1

SO

2

=1 a

tm

0.1

1E-6 1E-7

P

SO

PO2

1E-8

2

=0.1

atm

1E-9 1E-10 10

20

30

40

50

(Cu+Ni), wt%

60

70

80

0.01 30

40

50

60

70

(Cu+Ni), wt%

Fig. 3 Comparison of a the calculated S2 partial pressure (PS2) of Cu-Ni matte and b As distribution ratio between slag and matte with experimental measurements. Lines are calculated bu model. Symbols are experimental data [12]

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Viscocity Model The structure related viscosity model in MPE predicts viscosity of higher order slag systems by using binary parameters only. The model has been validated using published data in the higher order systems of slags containing oxide components of SiO2, Al2O3, Fe2O3, CaO, MgO, MnO, FeO, PbO, NiO, Cu2O, ZnO, CoO, CrO, Cr2O3 and TiO2. Very close agreement has been achieved over several orders of magnitude between model predictions and available viscosity measurements for a number of industrial slags. Details of the formulation of the model and most of the validation results can be found in Refs. [14–16]. Viscosity of solid-containing slags with up to 0.33 mass fraction of solid phases can also be estimated in the MPE package by using a modified Einstein-Roscoe equation [17].

Operation of Mpe The MPE package has a standalone user friendly Windows-based interface and a short training period is required for operating it. In addition to that, MPE can also be driven by Excel worksheet. The input data can be loaded in Excel worksheet, and the specified results can be displayed in worksheet. This feature allows MPE to do multi-stages flowsheet modelling by coupling the output of previous stage with the input of the next stage. Through an excel worksheet, MPE can exchange data with other software during a complex process modelling. In addition, large data sets, such as plant data, can be automatically processed batch wise by MPE once the input conditions are set up in Excel. The results can then be meaningfully statistically analysed and examined for significant trends. The plotting function in Excel also makes it convenient to visualize the variation trend of the results.

Applications As Deportment in Nickel Smelting Nickel sulpide concentrate smelting normally consists of smelting and converting stages. The nickel-copper concentrate, also containing minor amounts of cobalt and platinum group metals (PGMs) can be smelted to low grade matte during smelting stage which could occur in a flash smelter or electric furnace. The low grade matte is then normally transfered to a Pierce Smith (PS) converter to remove further Fe to produce low Fe matte for refining. Commercial nickel flash smelting was modelled based on the operational data listed in Table 1. The modelling results are compared with the plant observed

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Table 1 Operating conditions of the nickel outokumpu flash smelting [18] Feed chemistry, wt%

Ni

Cu

Co

Fe

S

SiO2

MgO

CaO

Al2O3

As

9.68 0.65 0.29 33.09 22.88 19.53 3.64 0.73 2.7 0.12 Charges, kg/h 76,242; Temperature, C 1350; O2-enriched air, Nm3/h 75,484 O2 pct in air 23.8 Oxygen efficiency, Pct 99.4 Heavy oil, kg/h 1,900

values in Table 2. It shows that there is general good agreement between the modelling results and the plant data. The calculated cobalt deportments to the slag and matte are close to the plant observed values. The modelling results show that As distribution ratio between slag and matte is about 0.017, which suggests that the slag removal of As is very small. The As deportment among gas, slag and matte phases during flash smelting under varied conditions such as matte grade, Ni/Cu ratio in feed material and CaO/ SiO2 ratio in the slag have also been simulated using MPE. Arsenic was evaporated into the gas phase as AsS, AsO and As2 etc. As shown in Fig. 4a, the dominant As gas specie is AsS at low matte grade region due to the high PS2. Further increase of matte grade leads to the increase of PO2 and decrease of PS2, hence the increase of the AsO pressure and decrease of the AsS pressure. With high grade matte, where PAsO is higher than PAsS, the AsO become the dominant As gas specie. Figure 4b shows that that at low matte grades, about 75 pct of the As reports to the gas phase with the balance being predominantly contained in the matte. However, as the matte grade increases to about 70% Ni, the relative portion of As that reports to the gas phase decreases to less than 25 pct with a corresponding increases in As deportment to the matte phases. This apparent effect of matte grade is caused by multiple contributing factors that also are dependent on matte grade. Similar to the copper matte, the activity coefficient of As in the Ni-Cu matte decreases with an increasing matte grade; thus, higher grade mattes have a higher affinity for arsenic and hence reduce the activity and vapor pressure of As. As discussed above, the S2 partial pressure in mattes also decreases with increasing matte grade, thus causing lower degree of volatilization of As as AsS at higher matte grades. It is also shown that the As deportment into the slag phase is low across the whole range of matte grade region. The As deportment to slag only contains a maximum value of 13 wt% at matte grade of 79 where the PO2 in slag reaches 2*10−7 atm. Figure 4c shows that the impact of Ni/Cu ratio in the feed on the As deportment at the fixed matte grade of 50. Increase of Ni/Cu ratio in the feed leads to the increases of As deportment to the matte. This is because the Ni has higher affinity for As than Cu, therefore, high Ni/Cu ratio in matte results in the more As been locked in matte during smelting. This results were supported by the findings that As distribution ratio between slag and nickel matte is much lower than that of copper matte [12].

slag

matte

observed calculated observed calculated

wt% wt% wt% wt%

18440 16300 49480 48900

Mass, kg/h 45.39 44.9 0.45 0.15

Ni 3.34 3.05 0.08 0.06

Cu 0.76 0.86 0.14 0.17

Co 22.68 22.3 44.73 44.22

Fe

Table 2 The comparison of the calculated slag and matte chemistry with the plant data 25.05 27.7 0.70 0.90

S 0.97 0 30.43 30.4

SiO2

0.40 0 5.38 5.67

MgO

0.56 0 1.17 1.13

CaO

As – 0.234 – 0.004

Al2O3 – 0 – 3.26

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Application of MPE Model to Nickel Smelting

(a)

347

(b) 100

1E-3

AsS

75

1E-4

As Distribution(%)

As gas species Partial Pressure, atm

o

T=1350 C silica saturation PSO2=0.1 atm

AsO

1E-5 As2

1E-6

o

T=1350 C silica saturation PSO2=0.1 atm

Slag

50

25 Matte

W Ni/W Cu=14

1E-7

30

40

50

60

0 30

70

(Cu+Ni), wt%

(c)

50

60

70

(Cu+Ni), wt% 100

o

T=1350 C silica saturation PSO2=0.1 atm

o

T=1350 C silica saturation PSO2=0.1 atm

Gas

Ni+Cu=50%

75

As Distribution(%)

75

40

(d)

100

As Distribution(%)

Gas

W Ni/W Cu=14

50

25

Gas

Ni+Cu=50%

50

Slag

25

g

Sla

Matte

0

Matte

0

1

10

Ni/Cu in concentrate, wt%/wt%

0.1

1

CaO/SiO 2, wt%/wt%

Fig. 4 The calculated a the partial pressure of As gas species, b As deportment versus matte garde, c As deportment versus Ni/Cu ratio and d As deportment versus CaO/SiO2 ratio in slag at 1350 °C

It has been demonstrated in flash copper converting that calcium ferrite slag is more effective in As removal due to its high As capacities than iron silicate slag [3]. The investigations by Font et al. [19, 20] show that the As distribution between calcium ferrite slag and nickel matte is about one magnitude higher than that of iron silicate slag, which suggests that adding more CaO into the slag could improve the As slag removal significantly. However the modelling results in Fig. 4d shows that the As deportment in slag only increases slightly with increasing CaO/SiO2 ratio in slag.

Properties of Nickel Smelting Slag Slag liquidus and viscosity are an important parameters in nickel smelting and converting processes. They affect the effective separation of the molten matte and

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slag, slag foaming behaviour and tapping of slag from furnaces. Slag viscosity also impacts mass transfer and hence could influence departure from equilibrium. Therefore, development of a thorough understanding of the relationship between the physico-chemical properties of the nickel smelting slag and its chemistry is expected to assist plant operations. It is thus of interest to estimate the effect of slag chemistry such as Fe/SiO2 ratios and MgO content on the slag liquidus and viscosities using the MPE model. As shown in Fig. 5a, the slag liquidus temperature goes through a minimum value as the Fe/SiO2 ratio in slag increases. Depending on the MgO level, the Fe/ SiO2 ratio of the slag with minimum liquidus varies. The primary solid phase is SiO2 at Fe/SiO2 ratio below than the minimum liquidus and olivine solid phase after the minimum point. The calculated viscosities of the slag at melting temperature was shown in Fig. 5b. Contrast to the slag liquidus, the slag viscosity goes through a maximum value as the Fe/SiO2 ratio in the slag increases. The initial changes of viscosity in SiO2 saturation region is due to the rapid decreases of liquidus temperature. In the olivine saturation region, increases of Fe/SiO2 ratio leads to the breakdown of the long silica chain formed in slag, hence reduce the slag viscosity. The results suggest that the optimum Fe/SiO2 ratio in slag should be determined by considering the combined effects of slag liquidus and viscosity on smelting performance. A batch of commercial nickel flash smelting data of Smelter A were analyzed by using MPE. As shown in Fig. 6, the liqudus of the slag falls in the range of 1280–1350 °C. The primary solid phase is olivine. The slag viscosities were calculated at the plant measured slag temperature. The slag viscosities are in the range of 0.1–0.2 Pa.s. The few outliers with high viscosity are because the slag temperature is below the liquidus, so the solid olivine was suspended in the liquid slag affecting the viscosities.

(a) 1400

(b) 1.0 CaO=1.13 wt% Al2O3=3.26 wt%

Mg

w O=9

CaO=1.13 wt% Al2O3=3.26 wt%

t%

1300

MgO

=6 w

Viscosiy, Pa.s

o

0.8

Olivine sat.

t%

1250

0.6

0.4 M

Slag liquidus, C

1350

=3 w

t%

t%

w

w

t%

=6

=9

0.2

=3 w

gO

t%

gO

MgO

SiO2 sat.

gO

M

M

1200

1150

0.0 0.6

0.8

1.0

1.2

1.4

Fe/SiO2, wt%/wt%

1.6

1.8

2.0

0.6

0.8

1.0

1.2

1.4

1.6

Fe/SiO2, wt%/wt%

Fig. 5 Impact of Fe/SiO2 ratio and MgO on the slag a liquidus and b viscosities

1.8

2.0

Application of MPE Model to Nickel Smelting

349

(a) o

Temperature, C

1400 1350 1300 1250 1200

Viscosity, Pa.s

(b)

0.6 0.4 0.2 0.0 1.00

1.05

1.10

1.15

1.20

1.25

1.30

1.35

1.40

Fe/SiO2, wt%/wt%

Fig. 6 Calculated a liquidus and b viscosities of commercial nickel smelting slags

Summary Thermodynamic models and databases have been developed for slag, matte and alloy phases, in which the behavior of minor elements usually present in base metal concentrates such as As, Bi, Sb, Pb, Se, Te, Sn, Co and Zn are covered. These databases have been developed through critical evaluation of various available thermodynamic and phase diagram data. Examples on model and database validation against the literature data on Ni and S solubility in slag, Ni and As distribution between slag and alloy and the As distribution between slag and Ni-Cu matte were presented. Application of MPE in modelling the commercial nickel flash smelting process shows that the calculated mass balance and the deportment of major and minor elements among phases are in reasonably good agreement with the plant observed values. The As removal into the slag is negligibly small during nickel smelting due to the lower As distribution ratio between slag and nickel matte. The As removal by evaporation decreases with increasing Ni/Cu ratio in feed. Increases of CaO/SiO2 ratio in slag only increase the As deportment to slag slightly. The correlations between the slag chemistry such as Fe/SiO2 ratio and MgO level and the slag liquidus and viscosities were modelled. The increases of Fe/SiO2 ratio leads to the increase of slag liquidus and decrease of slag viscosity in the region where olivine will be the primary saturated solid phase. At the fixed Fe/SiO2 ratio, the slag liquidus increases with increasing MgO content in slag. Analysis of a batch commercial nickel smelting slag data shows that the liqudus of the slag falls in the range of 1280 to 1350 °C and the slag viscosities are in the range of 0.1–0.2.

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Acknowledgements Financial support from CSIRO Minerals Resources is gratefully acknowledged.

References 1. Zhang L, Jahanshahi S, Sun S, Chen C, Bourke B, Wright S, Somerville M (2002) CSIRO’s multi-phase reaction model and its applications in industry. JOM 54(1):51–56 2. Chen C et al (2006) Thermodynamic modelling of minor elements in copper smelting processes. In: Paper presented at the Sohn international symposium. San Diego, California, 27–31 August 2006 3. Chen C, Zhang L, Jahanshahi S (2010) Thermodynamic modeling arsenic in copper smelting processes. Metall Mater Trans B 41B:1175–1185 4. Muggianu YM, Gambino M, Bros JP (1975) Enthalpies of formation of liquid alloys bismuth gallium tin at 723 K choice of an analytical representation of integral and partial thermodynamic functions of mixing for this ternary system. J Chim Phys Phys- Chim Biol 72:83–88 5. Schuermann E, Schulz E (1971) Liquidus and solidus curves of the systems copper-manganese and copper-nickel. Z Met 62:758–762 6. Feest EA, Doherty RD (1971) Copper-nickel equilibrium phase diagram. J Inst Met 99:102– 103 7. Predel B, Mohs R (1971) Thermodynamic study of molten nickel-copper alloys. Arch Eisenhut 42:575–579 8. Kulkarni AD, Johnson RE (1973) Thermodynamic studies of liquid copper alloys by electromotive force method: part II. The Cu-Ni-O and Cu-Ni systems. Metall Trans 4:1723– 1727 9. Gaye H et al (1984) Modelling of the thermodynamic properties of complex metallurgical slags. In: Paper presented at the second international symposiumon metallurgical slags and fluxes. Lake Tahoe, Nevada, 11–14 November 1984 10. Takeda Y, Ishiwata S, Yazawa A (1983) Distribution equilibria of minor elements between liquid copper and calcium ferrite slag. Trans Jap Inst Met 24(7):518–528 11. Pagador RU, Hino M, Itagaki K (1997) Solubility of Ni, Cu and minor elements in FeOx-SiO2-MgO slag equilibrating with nickel alloy. In: Paper presented at the fifth molten slags, fluxes and salts, Sydney, Australia, 5–8 Jan 1997 12. Font JM, Hino M, Itagaki K (1999) Phase equilibrium and minor elements distribution between iron silicate base slag and nickel-copper-iron matte at 1573 K under high partial pressure of SO2. Mater Trans, JIM 40(1):20–26 13. Kongoli F, Dessureault Y, Pelton AD (1998) Thermodynamic modelling of liquid Fe-Ni-Cu-Co-S mattes. Metall Mater Trans B 29B:591–601 14. Zhang L, Jahanshahi S (1998) Review and modeling of viscosity of silicate melts: Part I. Viscosity of binary and ternary silicates containing CaO, MgO, and MnO. Metall Mater Trans B 29B:177–186 15. Zhang L, Jahanshahi S (1998) Review and modeling of viscosity of silicate melts: Part II. Viscosity of melts containing iron oxide in the CaO-MgO-MnO-FeO-Fe2O3-SiO2 system. Metall Mater Trans B 29B:187–195 16. Zhang L, Jahanshahi S (2001) Modelling viscosity of alumina-containing silicate melts, S. Scan J Metall 30:364–369 17. Zhang L, Jahanshahi S, Sun S, Lim M, Bourke B, Wright S, Somerville M (2001) Development and application of models for pyrometallurgical processes. Mater Forum 25:136–153

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18. Tan P, Neuschutz D (2001) A thermodynamic model of nickel smelting and direct high-grade nickel matte smelting processes: part I. model development and validation. Metall Mater Trans B 32B:341–351 19. Font JM, Hino M, Itagaki K (1998) Minor elements distribution between iron-silicate base slag and Ni3S2-FeS matte under high partial pressures of SO2. Mater Trans, JIM 39(8):834–840 20. Font JM, Hino M, Itagaki K (2000) Phase equilibrium and minor elements distribution between Ni3S2-FeS matte and calcium ferrite slag under high partial pressures of SO2. Metall Mater Trans B 31B:1231–1239

Practice on Exploration of Oxygen-Enriched Converting Industrial Production by Kaldo Furnace Zhihua Wang

Abstract Through metallurgical calculation and process analysis based on metallurgical theory, it is aimed to raise the concentration of smelting oxygen and improve the update speed of reaction interface, accelerate the reaction and achieve a high-intensity smelting at the stage of converting. The PM plant has launched a theoretical research on model of high-intensity converting aiming at increasing the oxygen concentration of converting, carried out transformation of Kaldo furnace converting system, performed a series of industrial trial operations, conducted multiple rounds of tracking test and data comparison in reducing the furnace cycle, lowering the unit consumption of fuel and stabilizing the furnace lining life. The move has practically tested and verified the correctness of high-intensity smelting theory.



Keywords Kaldo furnace Oxygen-enriched converting Oxygen concentration Furnace cycle Furnace lining life





Outline of Kaldo Furnace Origin Kaldo furnace, also known as oblique oxygen blown converter, was successfully tested by B.Kalling in Sweden and put into operation at Domnavet Plant, so it was named by combining the first syllables of Kalling and Domnavet. Later, it was retrofitted by adding facilities for blowing oxygen, oxygen-fuel or other gas so as to control the temperature and atmosphere inside the furnace and enable it to be used for both converting of exothermic reaction and smelting and refining of endothermic reaction. Such a converter is called as top-blown rotary converter (TBRC for abbreviation). Z. Wang (✉) Jinlong Copper Co Ltd, Tongling 244021, Anhui, China e-mail: [email protected] © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_28

353

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Z. Wang

Basic Structure The furnace is a rotary converter with air blowing from the top, which consists of rotating system, tilting system, lance air supply and oxygen supply system (as shown in Fig. 1). The rotating system of the furnace is composed of furnace body, rolling ring, thrust roller, pressure roller and furnace body rotating mechanism. The system makes the furnace body rotate around its own axis by 360° at a rate of 1 ∼ 20 rpm. The tilting system of the furnace body which tilts around horizontal axis comprises trunnion ring and furnace body tilting mechanism. It can tilt forward or backward by 360° at a rate of 0.1∼1 r/min and stop at any angle position for accomplishing charging, smelting, discharging and other operations. The lance and gas hood are integrated for conducting fuel delivery, melt agitation and smelting gas exhausting.

Operating Features Due to adoption of top-blown and rotary furnace vessel, Kaldo furnace features full agitation, accelerating the multi-phase reaction of gas, liquid and solid materials and possessing favorable dynamic conditions in mass transfer and heat transfer. The rotation of furnace vessel allows uniform heating and even corrosion, conducing the

Fig. 1 Structure composition of Kaldo furnace

Practice on Exploration of Oxygen-Enriched Converting …

355

extension of furnace lining life. The use of oxygen for smelting and converting in the same furnace has strengthened the smelting process, shortened the process, increased the concentration of metal oxide (MO) in flue gas and facilitated the recovery.

Production Practice of Normal Air Converting by Kaldo Furnace Selection of Copper Anode Slime Treatment Process Along with rapid development of Chinese non-ferrous metals industry, anode slime has increased significantly. The metal components of anode slime from different sources are increasingly complex, the domestically original semi-wet process and flotation-smelting combined process have been severely restricted in application. In this case, TNMG, after rounds of demonstration, took the lead in building a project of 4000 t/a anode slime resource comprehensive utilization by applying Kaldo furnace pyrometallurgical process in China, which was put into operation successfully in 2009. The production practice for several years has proved that, compared with other processes at home, Kaldo furnace pyrometallurgical process has advantages of less equipment, short process flow, short product cycle, high recovery rate and good environmental protection effect etc.

Production Practice of Normal Air Converting Since it’s put into operation, the converting stage of Kaldo furnace pyrometallurgical process has been adopting 0.2–0.4 Mpa normal oil-free compressed air as a medium of oxidation reaction. During the initial stage of production, through continuous exploration and improvement, certain operating experience has been accumulated, indexes have been optimized and breakthrough has been made in some key indexes, i.e. shorter cycle of each heat, less specific energy consumption and lower silver content in slag. However, along with output increase and introduction of cost assessment, there exist some problems with normal air converting. (1) The compressed air is characterized by low oxygen content, inadequate oxygen concentration for actual reaction, long reaction process and low oxygen utilization. (2) A great number of air entering the furnace has caused lower furnace temperature, which requires to consume a lot of fuel for maintaining the temperature. (3) The longer converting duration has resulted in longer cycle of each heat and restricted the possibility of increasing the production capacity.

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Z. Wang

Theoretical Research on High-Intensity Smelting Process of Kaldo Furnace Establishment of Mathematical Model For the purpose to solve the problems of smelting caused by normal air, the PM Plant joined hands with a university to establish a metallurgical reaction kinetic model and mathematical model for stress analysis. The liquid level stress analysis is as shown in Fig. 2. A converting reaction rate formula has been obtained according to derivation process: ( 3 2 υ = f ðTÞn L CSe L ≤ L0 qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ð1Þ a ⋅ P0 sin β L0 = ρg sin α In the formula, υ— f ðTÞ— Lo — L— Cse — n—

Reaction rate inside the furnace at the stage of converting; Viscosity function related to temperature; Critical distance between lance and reaction liquid level; Operating distance between lance and reaction liquid surface; Se concentration inside reaction melt; rotating speed of top-blown furnace

Verification and Analysis of Factors For verifying the proportional relation between reaction rate and flow of compressed air, Based on the actual data of production, under the temperature of Fig. 2 Schematic diagram for lance operating during the converting operation

Practice on Exploration of Oxygen-Enriched Converting …

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Table 1 Average value data of converting and deselenization duration (h) under different temperature and different compressed air flow Compressed air flow/Nm3/h

600

1000

1200

1400

Temperature (1373 K) Temperature (1473 K) Temperature (1573 K)

9.8 7.5 5.5

7.6 3.8 3.5

5 2.5 1.8

7.5 2.4 2.0

1373 K, 1473 K and 1573 K respectively, constant speed of 10 rpm, cycle time under different flows is verified by adjusting the flow of compressed air. The average value data of converting and deselenization duration are obtained as shown in Table 1. The relative rate curves of converting reactions under different temperature and different compressed air flow are plotted according to the formula (1) as shown in Fig. 3. Conclusion: At the stage of noble lead converting, a lot of oxygen needs to be consumed. The research on mathematical model of high-intensity smelting indicates that reaction rate of converting/refining is directly proportional to pressure and flow of converting compressed air and corresponding to optimal operating height of lance under specific pressure. By controlling the flow and oxygen concentration of compressed air entering the converting lance, on the one hand, it provides oxygen required for impurities’ oxidation reaction with certain excess in order to promote the thermodynamic reaction. On the other hand, suitable amount of air will enter the nozzle, with high linear speed of gas at the outlet of nozzle, which may reach or exceed acoustic velocity. The gas is injected at a continuous and stable stream state to blow off the slag on the melt surface, expose fresh reaction interface and form dense bubbles for violent chemical reaction with the melt in a bid to achieve high-intensity smelting. Based on the theoretical research of mathematical model for Kaldo furnace converting, the PM plant set about adopting industrial test method to further verifying the compliance of high-intensity smelting theory. Fig. 3 Schematic diagram of relative reaction rate between compressed air and converting

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Practice of Oxygen-Enriched Converting Used in Industrial Production Design, Installation and Commissioning The transformation project of oxygen-enriched converting by Kaldo furnace was initiated in 2013, with maximum oxygen supply capacity of 1,000 Nm3/h, maximum operating pressure of 0.6 Mpa, maximum oxygen supply capacity of 600 Nm3/h for converting valve block, operating pressure range of 0.25 Mpa ∼ 0.55 Mpa for engineering design. The set target oxygen concentration is achieved by choosing converting nozzles of different rated flow and setting fixed compressed air flow and oxygen flow regulating valve. The maximum regulating range of oxygen concentration for lance converting is 21% ∼ 95%. The range of transformation is as shown in Fig. 4 and the control loop is as shown in Fig. 5.

Fig. 4 Diagram of newly-added converting oxygen supply pipeline

Fig. 5 Control diagram of oxygen-enriched converting

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In May 2015, Tongling Nonferrous Metals Design Institute completed the design of working drawings for newly-added oxygen pipeline. In July 2015 when the working drawings passed the drawing review, the oxygen-enriched converting pipeline stepped into installation and commissioning. In June 2016, the project officially obtained the qualification test report of pressure pipeline for the newly-added pipes and conducted the final inspection and linkage test before trial production.

Industrial Trial Production Preliminary Trial Production Based on Step of Oxygen Concentration For conducting industrial trial production smoothly, a trial production leading group was set up for repeated deliberation of air-oxygen proportion and operating procedures and finally selecting of trial production program based on three steps of 25%, 27% and 30% oxygen concentration. The comparison of data based on three steps of oxygen concentration is as follows (Tables 2, 3 and 4): As shown in the data from the above tests based on three steps of oxygen concentration, under the case of oxygen-enriched converting based on 25% and 27% oxygen concentration for treatment of 15–16 ton decopperized anode slime, the converting duration is stabilized at 5–6 h, showing that the converting duration has not been shortened notably and the oxygen-enriched converting effect is unapparent; under the case of oxygen-enriched converting based on 30% oxygen concentration for treatment of the same tonnage decopperized anode slime, the converting duration is reduced by nearly 2 h per heat on average.

Trial Production Based on Step of Dore Metal Output For further verifying the oxygen-enriched converting test result under the oxygen concentration of 30%, the group decided to continue the test based on step of dore metal output. The summarized data of the test are as shown in Tables 5 and 6: (1) For oxygen-enriched converting based on oxygen concentration of 30%, the duration can be reduced by at least 1.5 h per heat, with reduced cycle of 1.5–3 h per heat on average. (2) For oxygen-enriched converting based on oxygen concentration of 30%, the natural gas consumption can be reduced by at least 231 Nm3 per heat, with reduced natural gas consumption of 231–329 Nm3 per heat on average.

299 286 293 360 274 295 304 302 301.6

25 26 27 28 29 31 32 33 Average

25 25 25 25 25 25 25 25

Converting duration (min)

Oxygen-enriched converting Heat Mixed oxygen no. concentration (%) 15004 14997 15137 15009 15127 15001 14900 15035 15026.3

Charging amount of anode slime (kg) 32 33 34 35 36 37 38 39 Average

21 21 21 21 21 21 21 21

Air converting Heat Air oxygen No concentration (%) 403 283 334 341 336 298 393 311 337.4

Converting duration (min)

Table 2 Comparison of data between 25% oxygen concentration oxygen-enriched converting and air converting

15548 15572 15626 15914 16150 16196 16191 16078 15909.4

Charging amount of anode slime (kg)

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Table 3 Main data for oxygen-enriched converting of 27% oxygen concentration Heat no.

Mixed oxygen concentration (%)

Converting duration (min)

Charging amount of anode slime (kg)

35 36 37 38 Average

27 27 27 27

330 258 328 342 314.5

15869 16063 16003 16036 15992.8

Table 4 Main data for oxygen-enriched converting of 30% oxygen concentration Heat no.

Mixed oxygen concentration (%)

Converting duration (min)

Charging amount of anode slime (kg)

40 41 42 43 44 45 Average

30 30 30 30 30 30

187 184 241 158 248 186 200.7

15978 16011 16034 15920 16047 16015 16000.8

Table 5 Statistical analysis on corresponding cycle and fuel consumption of different dore metal outputs Class of dore metal output Class Class Class Class Class

of of of of of

1400 1500 1600 1700 1800

kg kg kg kg kg

Shortened cycle per heat on average (h)

Reduced fuel consumption per heat on average (Nm3)

3 2.2 2.5 3.1 1.5

316.6 258.6 252.1 328.8 231

(3) The oxygen-enriched converting started to continue the test on October 2, 2016. After October, the unit consumption of oxygen in the pipeline rose slightly. The average unit consumption of oxygen in January-September is 261.48 m3/t slime. From October 2016 to January 2017, the average unit consumption of oxygen is 288.27 m3/t slime, with average increase of 26.79 m3/t slime.

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Table 6 Data of oxygen consumption in pipeline in January 2016 and January 2017 Month

Total consumption (m3)

Unit consumption (m3/t slime)

Jan. 2016 Feb. 2016 Mar. 2016 Apr. 2016 May. 2016 Jun. 2016 Jul. 2016 Aug. 2016 Sep. 2016 Oct. 2016 Nov. 2016 Dec. 2016 Jan. 2017

119889 151140 142182 157297 118889 141832 119470 152916 99394 154800 187700

294.950 272.193 247.918 274.274 237.047 262.813 224.387 276.574 263.166 270.427 317.581 291.892 273.179

165920

Impact upon Lining Consumption Rate After Oxygen-Enriched Converting During the trial production of oxygen-enriched converting, tracking record was also made for lining consumption of Kaldo furnace (Table 7). (1) It shows that, after oxygen-enriched converting is adopted, the lining consumption rate of the first 45th heats during the 23–27th campaigns has not been increased obviously and remains within the expected range. (2) Judging from the 23–25th campaigns in which air converting is used, the lining consumption of the middle and lower parts is more obvious than that of the middle and upper parts. By contrast, for the 23–27th campaigns with oxygen-enriched converting adopted, the lining consumption of the middle and lower parts is obviously lower than that of the the middle and upper parts. Table 7 Comparison of input, output and hearth consumption for thee 23th–27th heats and the first 45th heats Heat no.

23

24

25

26

27

Furnace campaign.heat Consumption of hearth depth/ maximum.mm Consumption of middle and upper linings/maximum.mm Consumption of middle and lower linings/maximum.mm Total input of material. t Dore metal output. t Effective operating duration. h

92 120

85 100

86 94

88 80

– 60

365

335

415

355

385

395

365

495

315

335

1047.752 71.729 1224.6

987.335 78.22 1212.6

1049.285 73.444 1153.7

902.457 69.536 1035.3

975.945 67.263 1076

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Economic Result According to the conclusion of the above analysis, we made an analysis on the economic benefit made by oxygen-enriched converting of Kaldo furnace under the prerequisite of no impact upon the consumption rate of Kaldo furnace lining’s working layers. (1) Benefit made by shortened cycles Based on the above conclusion, the adoption of oxygen-enriched converting can reduce 1.5–3 h per heat on average. If 2.25 h is taken as an intermediate value, 272 heats and 271 heats were produced in 2015 and 2016 respectively. If 270 heats are based on for each year, the total hours saved will be: 270 heats × 2.25 h = 607.5 h. If the calculation is based on 25 h for each heat, 1 h interval between heats and 3.7 h for clearing of each heat, it is expected to increase the heats of operation per year: 607.5 h ÷ (25 + 1 + 3.7)h/heat = 20.5 heats. If 15t decopperized anode slime is treated for each heat, it is expected to increase the treatment of decopperized anode slime per year: 20.5 heats × 15t = 307.5 tons; The treatment of anode slime increased per year: 307.5 tons (decopperized anode slime) × 0.65% (decoppering slag rate) = 473 tons; The treatment of anode slime increased 473t ÷ 5,900t ≈ 8% indicates that the reduction of fixed cost for treatment of copper anode slime is around 8%. The treatment cost of copper anode slime for the PM plant is RMB12,000 per ton, including controllable cost of around RMB7,500 per ton and fixed cost of around RMB4,500 per ton. After implementation of oxygen-enriched converting, it is expected to save the cost: RMB4,500 per ton × 8% × (5,900 + 473) tons = RMB2.29428 million. (2) Cost saved for reduction of natural gas consumption The conclusion of analysis shows that natural gas saved per heat can reach 231–329 Nm3. If 280 Nm3 is taken as an intermediate value, the natural gas can be saved by 270 heats × 280 Nm3 = 75600 Nm3 per year. If unit price of natural gas is RMB2.59 per ton, the natural gas can be saved by 756 × 2.59 = RMB195,800 per year. (3) Cost increased for oxygen consumption of pipeline According to the statistical data for oxygen-enriched converting of the 26th campaign, the adoption of oxygen-enriched converting will increase the oxygen consumption of 537.8 Nm3 per heat on average, the annual oxygen consumption increased will be 537.8 Nm3 × 290 heats = 155,962 Nm3. It it is calculated based on RMB0.51/Nm3 unit price of pipeline oxygen, the annual production cost increased is RMB 0.51/Nm3 × 155,962 Nm3 = RMB79,500.

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Investment Cost The expenses for early-stage preparation and design of oxygen-enriched converting were RMB20,000, with RMB110,000 used for equipment purchase and RMB40,000 for installation and commissioning, totaling RMB170,000. Benefit in the first year of oxygen-enriched converting: RMB2.29428 million + RMB195,800 - RMB79,500 - RMB170,000 = RMB2.24058 million. A benefit of RMB2.24058 million can be made per year. (No consideration has been given to the change of lining loss and equipment maintenance cost.

Summary After trial operation of a period, the oxygen-enriched converting operation has basically grasped the application regularities and optimal operating conditions. Presently, it has been formally put into practical operation. (1) The adoption of oxygen-enriched converting allows to save at least 1.5 h of cycle per heat, increase the treatment of copper anode slime by 473 t and make an economic benefit of RMB2.41 million due to reduction of operation and maintenance costs. (2) The successful utilization of oxygen-enriched converting has provided a reference for TNMG’s Phase II precious metal smelting project in equipment selection, localized renovation of Kaldo furnace as well as other similar plants. (3) During the oxygen-enriched converting operation, some metallurgical technical indexes need to be improved, therefore, further tracking analysis is required to be made in the subsequent production practice. (4) After oxygen-enriched converting, the consumption of working layers in the middle and upper parts of the lining are faster than those in the middle and lower parts. It is may be caused by oxygen-enriched converting or high slag line, which is subject to further practical illustration.

Ust-Kamenogorsk Metallurgical Complex: A Silent Achiever Alistair Burrows and Turarbek Azekenov

Abstract Situated in the oblast of East Kazakhstan, the Ust-Kamenogorsk Metallurgical Complex has spent the past twenty years renovating and remodelling itself as a modern polymetallic sulphide smelting facility, and a regional centre of excellence for custom smelting. Along the way there have been changes to the company structure, introduction of new technologies, investments in expanded capacity, environmental improvements, and addition of new metal products and by-products to the site’s repertoire. These changes have been gradual and incremental, but taken together they represent a significant contribution to placing Kazakhstan’s sulphide smelting industry on a strong foundation for enduring success in the international custom smelting market. In achieving these changes, a workforce that was historically isolated from much of the world now has recognised expertise, internationally competitive skills, and confidence to embrace the future. Further improvements in energy efficiency, environmental compliance and polymetallic processing capabilities are challenges that UKMC stands ready to face. Keywords Copper



Zinc



Lead



Precious metals



Kazakhstan

Introduction Formerly part of the Soviet Union, the Republic of Kazakhstan has been an independent nation for more than 25 years. It is larger than Western Europe and has a wealth of mineral resources spread throughout the country. The capital of Eastern Kazakhstan oblast is Ust-Kamenogorsk, a city of about 350,000 people, which has been a hub for transporting and processing base metals concentrates for more than a century. For many years Ust-Kamenogorsk has also been home to a multi-smelter metallurgical complex specialising in the production of zinc, lead, copper, silver, gold, antimony, bismuth and various other by-products. The Ust-Kamenogorsk Metallurgical Complex (UKMC) A. Burrows (✉) ⋅ T. Azekenov Kazzinc Ltd., 1 Promyshlennaya Street, 070002 Ust-Kamenogorsk, Republic of Kazakhstan e-mail: [email protected] © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_29

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has an interesting history, a series of notable metallurgical achievements in the field of sulphide smelting and bright prospects for the future. The development of UKMC is best understood when set in its geographic, political and economic context. Both the advantages and the challenges UKMC face include features that are not commonplace in sulphide smelters elsewhere. This paper seeks to introduce the modern plant to a wider audience, so that its statistics and achievements might be better known and discussed.

History and Geography In the wide area of Central Asia there have been many rich deposits of non-ferrous and precious metals, known from the time of Imperial Russia. The Altai mountains were a particularly prospective area [1] and one rich mine, near today’s city of Ridder, has been in constant operation for more than 200 years. Ust-Kamenogorsk lies at the foothills of the Altai mountains, at the confluence of the Irtysh and Ulba Rivers. Road, rail and (in past times) river are the major transport routes for trade. Ust-Kamenogorsk was proclaimed a city 150 years ago, and its development has been closely linked to the mining and metallurgical industry ever since. In the early 20th century a port was built on the Irtysh River, a rail link was built to the mines of Ridder [2, 3] and Ust-Kamenogorsk became a transport hub for the mineral wealth of the region. The city was developed into a major metallurgical processing centre during the Soviet period. Non-ferrous metals, especially uranium, beryllium, titanium, magnesium, copper, lead, silver and zinc were produced in various parts of the city. Zinc production began in 1947, and lead production began in 1952. These two metals form the core of what is today’s UKMC, which was for many years the headquarters of zinc and lead metallurgical development within the Soviet Union. From its technical and operating expertise many processes were developed. Among the most famous is the KIVCET furnace for smelting lead concentrates, which has been adopted by smelters in Italy, Canada and China. In past decades some of Ust-Kamenogorsk’s achievements in metal processing have been well publicized internationally, but other achievements have not received similar attention. This may in part be due to Ust-Kamenogorsk’s post-war industrial history being associated closely with uranium refining. The city was therefore kept closed to outsiders, and the isolation of this time may have helped to foster some local trademark professional attitudes such as resilience, resourcefulness, self-sufficiency, along with a tiny bit of pride.

The Formation of “Kazzinc” The disintegration of the Soviet Union brought with it an upheaval of the socio-political system, which for several years had a negative impact on the mining and metallurgical production of lead, zinc and precious metals in the area. By 1996,

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the Kazakh government faced the fact that the nonferrous metallurgy enterprises it owned in East Kazakhstan were in decline [4]. There had been investment in local precious metals refining capacity, but in all other respects the plants of the region were deprived of new capital. With little investment for the past 10 years, it was no longer possible to maintain a cost-effective and coordinated operation of the mines, concentrators and smelters. As a result, the three main non-ferrous metal companies in the East Kazakhstan region, plus a local hydroelectric power plant, were merged into a single corporatized entity suitable for privatisation. These entities were—“Ust-Kamenogorsk Lead and Zinc Combinate”, “Leninogorsk Polymetallic Combinate” and “Zyrianovsk Lead Combinate”. The company “Kazzinc” was formed in January 1997, with Glencore International becoming the company’s main investor. Some issues facing the new company of Kazzinc included: • A need to improve environmental performance • A need to bring forward the development of new mines to replace diminishing reserves • An urgent need for “catch-up” maintenance to overcome recent neglect • A need to improve the economics of all fixed plant assets, by investment in modern technology Over the twenty years that followed the privatization, there has been a gradual capital investment of over USD$1.5 billion to progressively address these needs. The process has not been fast, but the transformation has been thorough, noteworthy and rewarding. The ongoing development of UKMC has been at the core of the transformation. Figure 1 shows a schematic representation of UKMC’s place within the broader Kazzinc group.

Fig. 1 Structure of Kazzinc

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Zinc As the company’s name implies, historically the major focus of Kazzinc Ltd has been the production of zinc. Construction of UKMC began during World War II using equipment evacuated from the Vladikavkaz zinc smelter in 1941 and continued after the war when equipment from Magdeburg zinc plant was brought from Germany as part of war reparations. A lack of design drawings severely complicated and delayed the reassembly and integration of the Magdeburg equipment, resulting in plant commissioning in 1947. Today’s production is fully integrated, with the vast majority of zinc sales being in the form of alloys or refined zinc ingots, and a small amount as zinc sulphate that is used as a flotation reagent at Kazzinc’s concentrators. In total, the company accounts for around 95% of Kazakhstan’s zinc production [5], with around 60% processed via UKMC. In the first 5 years following the formation of Kazzinc the investment at UKMC’s Zinc Plant amounted to tens of millions of dollars in like-for-like refurbishment of the dilapidated roasting and tankhouse equipment. The plant capacity is now approximately 190,000 tpa of zinc. The major concentrate supply to the Zinc Plant is a sulphide zinc concentrate (assaying approximately 50% Zn) originating from the Maleyevskiy mine. The Zinc Plant uses a conventional RLE production process, as shown in Fig. 2. There are some local idiosyncrasies. Concentrates are treated in oxygen-enriched fluid bed roasters. Although not widespread elsewhere, the use of oxygen-enrichment in zinc fluid bed roasters was commonplace among former Soviet countries. The oxygen concentration in roaster blast air may vary between 21–39%O2 and in practice it is rarely the heat balance that limits the Zinc Plant’s use of tonnage oxygen, but rather internal competition with hungry oxygen consumers in the Copper Plant and the Lead Plant.

Zinc concentrates

Zinc Dust Zinc drosses DMT 5138 Zn 87,6%

Concentrate Storage

Filtration Thickening

Calcine Leaching

Zinc saleable

Cementation

Spent electrolyte Zn 50 g/l Roasting section

electrolysis

Calcine

Zinc dust

Cadmium section

Cadmium sponge Copper Cement Roasting gases 70 000m3/h SO2 - 13,5 % vol.

To Sulphuric Acid Plant

To Slag Fumer

Neutral solution Zn 145 g/l

Cadmium metal

Ferrous solution Zn 75 g/l filtration

Zinc residue

Slag fumes and Waelz oxides leaching

Used as Reagent

Thickening Cementation

Waelz Kiln Waelz Oxide

Clinker

To Magnetic Separation

ZnO Fumes

From Lead Plant

Fig. 2 Schematic flowsheet of UKMC Zinc Plant

Lead residue To Lead Plant

Copper chloride residue

To Copper Plant

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UKMC has three fluid bed roasters, of which sometimes one and sometimes two are operational at any given time. Roaster off-gases are captured and converted into sulphuric acid in a contact acid plant. Zinc calcine from the roasting process is leached with spent electrolyte. Copper and cadmium are removed from the resultant zinc solution. Cadmium is refined for sale while the copper is removed as copper chloride, which becomes one of the feed materials for UKMC’s Copper Plant. Zinc residues are processed pyrometallurgically in Waelz kilns, producing Waelz oxide from which lead, chlorine, arsenic and antimony are removed prior to treatment in the calcine leaching circuit. The Waelz slag is further processed for recovery of useful carbon, and valuable copper and precious metals. These are examples of the integrated nature of UKMC processing plants. So while Kazzinc is a large integrated zinc producer, with operations spanning 6 towns in Kazakhstan, UKMC has become a metallurgical hub that links all of the operations and provides benefits that can only be achieved when treatment of by-products is holistically integrated into the company operations.

Maleyevskiy Mine Although this article is principally focused on sulphide smelting at UKMC, a brief mention of Maleyevskiy mine is necessary to appreciate the motivating force behind our site’s development of the last 20 years. The Maleyevsky mine is Kazzinc’s largest underground operation, commissioned by the company in June 2000. By the end of 2001 the mine had been expanded, bringing production up to 2.25 Mt/y of ore, with average ore grades of Zn 7.5%, Cu 2.3%, Pb 1.3%, Au 0.75 g/t and Ag 75 g/t [6]. The proven reserves at that time suggested approximately twenty years of mine life. Maleyevskiy mine accounts for the vast majority of UKMC’s zinc production and about 50% of UKMC’s copper production. Maleyevskiy lies 18 km east of the town of Zyrianovsk, so after the mine was commissioned an existing concentrator at Zyrianovsk was used to process the ore. Previous ores from Kazzinc’s mines were mainly Zn-Pb ores, but since its inception the Maleyevskiy mine has produced a Zn-Cu ore. Of recent times Maleyevskiy ore typically contains the following grades: 5.7% Zn, 1.9% Cu, 0.9% Pb, Au 0.56 g/t Au and 56 g/t Ag. Concentrates produced from Maleyevskiy ore have some characteristic features. The copper mineralisation is distributed between chalcopyrite: tennantite in the approximate mass ratio of 15:1, which limits the potential smelting outlets for the copper concentrate. In the years since 2001 an increasing dependence on Maleyevskiy mine has driven UKMC’s development to deviate from its history and strike out in a new direction: it has forced the company to become a copper producer, and it has forced the Lead Plant to change from a captive smelter to a custom smelter.

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Copper In 2005 Kazzinc instigated a project to build, from scratch, a copper smelter, another acid plant and a copper refinery to complement the existing processing plants within UKMC. Despite its rich smelting history, copper had not previously been among UKMC’s major products. The new plant would be akin to a greenfield project in scope, except that its location amid a major metallurgical complex had all of the complications associated with a brownfield expansion project. The concept of the new Copper Plant was that it should have a nominal production capacity of 70,000 tpa of cathode copper, be able to treat polymetallic copper concentrates and a range of by-products from zinc and lead refining, be tolerant of minor element fluctuations, and be readily expandable in the future. The flow sheet of the copper plant, shown in Fig. 3, has been described by others [7, 8]. It includes the following process stages: • smelting of copper concentrates and by-product materials in an ISASMELT™ furnace; • settling of matte and slag, and slag cleaning, in an electric furnace; • copper matte converting in Peirce-Smith converters; • fire refining of blister copper in rotary anode furnaces; and

Fig. 3 Flowsheet of UKMC Copper Plant

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• electrorefining of anodes on IsaProcess™ permanent cathode plates. Copper and gold concentrates, zinc plant residues and copper-bearing reverts are smelted in the ISASMELT™ furnace to produce a gold- and silver-bearing matte. Dust produced in the ISASMELT™ is leached to provide an arsenic bleed, and for recovery of lead. De-dusted gas is sent for sulphur dioxide capture at a new Acid Plant. Copper anodes, after fire-refining, are sent to the tankhouse for production of refined copper with purity predominantly corresponding to M00 K brands (equivalent to LME grade ‘A’). Gold- and silver-bearing copper anode slimes are treated at the Precious Metals Refinery. Kazzinc’s mined copper represents only 10% of Kazakhstan’s production [5]. With a surplus of concentrates and a deficit of smelters the UKMC Copper Plant was a necessary asset for an integrated producer to capture added value, but a Copper Plant with 70,000 tpa capacity for refined copper cathodes is remarkably small by modern standards. The plant is economically viable only because of several local anomalies such as: copper residues being available from zinc plants, copper dross being available from the lead smelter, cheap electrical power from the Bukhtarma hydroelectric power station being available for the copper refinery, and a tariff applying to the export of copper concentrates from the Republic of Kazakhstan. As the Copper Plant production grows and exceeds the output of Maleyevskiy mine, UKMC is looking farther afield at potential polymetallic copper concentrates that are available elsewhere in Central Asia, and which may provide an attractive feed for the smelter. The Copper Plant is gaining the flexibility required to treat various feed types in the future.

Lead Today lead from UKMC accounts for nearly 100% of the lead production in Kazakhstan [5]. For the first five years after the formation of Kazzinc, the principal task was to rationalize the lead production facilities. Crude lead production previously had been conducted at Ridder with pyrometallurgical refining centralized in UKMC. This was changed so that all lead production was centralized in UKMC, which had immediate environmental benefits. The atmospheric dispersion of SO2 was thus greatly reduced in Ridder. If UKMC is the “hub” of Kazzinc, then the Lead Plant is the “hub” of UKMC. The Lead Plant is responsible for providing the primary feed to the Precious Metals plant, and for accepting all manner of by-product materials that come from the Zinc Plant and Copper Plant. Schematically the relationship between the plants resembles the diagram in Fig. 4.

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Fig. 4 Relationship between Lead Plant and other plants of UKMC

The Lead Plant is a sprawling complex that has grown over many years. Construction of the Lead Plant began in January 1951. The first metal was produced on June 25, 1952. In January 1956 a slag-fuming plant was started up for recovery of zinc from the blast furnace slag. An updraught sinter plant, of Lurgi design, was installed in 1974 and most of the equipment in the extant sinter plant dates from that time. A KIVCET furnace with production capacity of 40,000 tpa crude lead bullion operated from 1987-1997, but since 1997 all lead production relied on the sinter plant—blast furnace production route, until in August 2012 an ISASMELT™ furnace was introduced as a part of the New Metallurgy Project [9, 10]. This process broadened the range of treated lead-bearing products compared with sintering. Ingots of lead-rich slag are added to the blast furnace, at grades which were considered unfeasible with sintered agglomerates. Complete sulphur removal prevents matte generation in the blast furnace and has resulted in further reduction of atmospheric dispersion of SO2 gases from the UKMC site. The Lead Plant modernization was conceived in 2005 at the same time as the Copper Plant construction project was instigated. What was not expected at that time, was that the lead ISASMELT™ furnace would go on to exceed its design capacity by 40% and contribute to a significant reduction in the costs of lead production at UKMC. This year’s 151,000 tonne of refined 99.99% Pb represents the highest production forecast in decades. It has resulted in a re-think for the operation of many downstream processes and equipment items. Kazzinc proceeded with the modernisation of its Lead Plant flowsheet assuming that the existing lead blast furnaces would be satisfactory for treating the slag produced in the lead ISASMELT™ plant, with little or no capital investment required. This assumption has since been validated. The instantaneous capacity of the ISASMELT™ furnace to generate slag is approximately twice the instantaneous capacity of a single Kazzinc blast furnace to smelt slag, so two blast furnaces are in use most of the time. The present flowsheet is represented schematically by Fig. 5.

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Fig. 5 Present Lead Plant flowsheet for UKMC

Precious Metals Within UKMC both the Copper Plant (anode slimes) and the Lead Plant (Parkes crust) have significant by-products of precious metals that need to be refined. Parkes crust from the Lead Plant is distilled electrothermically to produce a predominantly lead-silver alloy for cupellation. Dried anode slimes from the Copper Refinery is also added to the cupels. The doré metal produced in the cupels is spooned manually for casting into moulds, creating anodes ready for refining. Until 1993, the refining was done in Russia. Kazzinc’s precious metals refinery was the first in Kazakhstan [11]. It dates from the period, just after independence, when it was constructed for reasons of establishing national gold reserves in the newly-independent Republic of Kazakhstan. The refinery was developed by local experts using a unique design. It has a compact silver refining area where particularly large doré anodes are refined to produce silver cathode granules. This allows more precious metal production in a small plant footprint. The purity of gold and silver produced is “four nines”. The refinery was added to the LBMA’s Good Delivery List in 1995 under the “Deer” brand. The precious metals refinery has proven to be a versatile asset. In recent years Kazzinc has developed (dedicated) gold mining as a new business venture. The Altyntau Kokshetau mine (literally translated from Kazakh it is the “mountain of gold”) now produces a substantial portion of the company’s revenue. Doré metal

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from Altyntau Kokshetau is brought to UKMC for processing into refined gold and refined silver. The annual capacity of the precious metals refinery is approximately 45 tons of gold and 1000 tons of silver.

Sulphur An acid plant was already in use at UKMC at the time of Kazzinc’s formation, but it was connected only to the roasters of the Zinc Plant. Reducing sulphur dioxide gas emissions across its metallurgical operations was one of Kazzinc’s founding ambitions. As described above, the first step was to centralize lead smelting operations at UKMC. In 2002 this step was followed by constructing a “wet sulfuric acid” plant at UKMC using the Haldor Topsoe technology to harvest “strong gases” from the strand of the Lead Plant sinter machine. It has been in operation since 2004. In practice, the wet sulfuric acid plant achieves excellent sulphur capture (tail gas strength is typically around 500 ppm SO2). With the 2005 decision to construct a Copper Plant, a new sulphuric acid plant was required. A double-contact/double-absorption acid plant using MECS technology was installed. The design of both the acid plant and the gas cleaning system allowed for gases to be harvested from a single blowing PS Converter, a single Copper ISASMELT™ furnace, plus the possibility of some surplus gases from the Zinc Plant or the Lead Plant operations. This results in a complicated “spaghetti” of

Fig. 6 Arrangement of smelters and acid plants at UKMC

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large duct connections, as shown in Fig. 6, but it has added to the overall operating flexibility of UKMC because in the event of an Acid Plant breakdown there is the possibility by adjusting damper positions to choose, and change, which smelter is idle and which operational. UKMC continues on the path of incremental reductions in SO2 emissions, and in the past twenty years has succeeded in reducing by more than 50,000 tpa the SO2 that was previously released to the atmosphere, all while increasing the production of metals. UKMC is now in full compliance with local environmental standards.

People Kazzinc is Eastern Kazakhstan’s largest industrial producer, and employs about 22,000 people in mining, mineral processing, metallurgical smelting and refining operations, as well as in electricity generation and machinery construction. The company strives to develop and improve its application of mining and metallurgical industry technologies. Some of the technology has been developed in-house, some is sourced from Russian-speaking companies based within former Soviet countries, but a large degree of Kazzinc’s success may be attributed to its willingness to learn from, and import equipment and knowhow from, leading companies all around the world. Over a twenty year period there has been a continuous and swift pace of learning by the workforce. The change should not go unremarked, because it is almost certainly the strength that has ultimately given rise to UKMC’s relative success. In this twenty year period the workforce has in some places adopted mechanization, automation, sophisticated control systems and safer work culture that were not within the employees’ dreams at the time of Kazzinc’s formation in 1997. No part of the company is untouched by this culture of development, but UKMC is a particular area of focus owing to the complexity of its operations. To learn the operations of various new metallurgical technologies adopted in the past twenty years, UKMC has selected employees and organised for them to spend time training in Sweden, Australia, Zambia, China and many other countries. The training itself and the various working trips to learn from foreign specialists have broadened the ambitions of each trainee involved. This process has nearly come full circle, most recently a few foreign companies have expressed interest in sponsoring their employees to travel to the Republic of Kazakhstan to spend time learning the operation of metallurgical technologies in common use at UKMC [12]. This is a source of considerable pride for those who have travelled the whole journey from trainee to trainer. Approximately 5500 people are employed at UKMC in metallurgical processing and various support services. The challenge for the next decade will be to manage the attrition of experienced workers with a gradual infusion of young employees. There are ambitions to realise further improvements in labour productivity across the site.

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Production As a result of sustained investment in fixed plant assets, training and work practices, UKMC has increased its output of base and precious metals during the past twenty years. Aside from the zinc smelter, which is a mature plant with a relatively stable feed suite, the production increase has been “bumpy” rather than steady, as illustrated in Fig. 7. In particular there have been rises and falls in precious metals as different mine sources have become abundant and then scarce again. At present the Lead Plant has over twenty different external suppliers in addition to ten suppliers internal to Kazzinc. Fluctuations in the amounts from different suppliers can have a significant effect. Melal production from UKMC 2000

2001

2002

2003

2004

2005

2006

Metal Lead

mt

133 925

107 584

101 643

88 741

99 149

88 929

86 198

91 025

90 446

79 501

101 023

100 989

83 356

90 982

125 390

120 105

134 059

147 284

1 870 328

Metal Zinc

mt

145 077

156 515

165 176

172 117

175 722

178 528

180 313

184 491

189 630

190 382

191 148

190 024

190 781

190 134

194 261

193 750

194 407

201 212

3 283 666

Metal Copper (Cu mt blister+ Cu cathodes) Metal Gold Metal Silver

toz

5 866

7 002

231 213

267 095

toz 11 148 538 9 285 795

6 512

7 787

7 727

6 133

5 904

2007

2008

5 345

2009

4 601

2010

6 211

2011

5 327

226 346

205 335

221 609

171 524

124 251

125 311

135 777

185 189

294 771

5 842 190

5 928 055

5 027 105

5 503 479

4 029 398

3 766 532

3 499 363

2 826 489

3 544 518

2012

2013

2014

2015

2016

2017

16 426

52 061

58 827

58 186

62 249

64 790

62 914

398 010

547 577

633 113

610 947

566 529

621 163

575 981

6 775 603 18 072 766 17 859 008 25 090 207 30 059 877 27 213 579 21 945 304

Total

443 868 6 141 742 207 417 804

700 000

35 000 000

600 000

30 000 000

500 000

RH Axis

25 000 000

400 000

20 000 000

300 000

15 000 000

200 000

10 000 000

100 000

5 000 000

0

0 2000

2001

2002

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Metal Lead mt

2004

2005

2006

Metal Zinc mt

2007

2008

2009

2010

2011

Metal Copper (Cu blister+ Cu cathodes) mt

2012

2013

Metal Gold toz

2014

2015

2016

2017

Metal Silver toz

Fig. 7 UKMC metal production

Complementary with the five major production metals, a variety of impurity elements are processed through to final products. These products include small amounts of Selenium, Indium, Thallium, Mercury, Bismuth, Tellurium, and Cadmium as well as sodium antimonate. In some cases the equipment for production of these metals is an inheritance from the plant architecture as it existed many decades ago. However, UKMC foresees an ongoing role in processing most of these impurity elements through to final products. UKMC lies within the city boundaries of Ust-Kamenogorsk and residential buildings exist within 300 m of the smelter production areas. UKMC has a “zero emission” rule for liquid water, and disposal of any impurity elements other than by vitrification in inert iron silicate slag would likely not receive regulatory approval.

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Power and Transport The availability of, and prices of, electrical power and bulk transport are respectively the largest structural advantage and disadvantage of UKMC’s future operation. Much planning is required to formulate annual production budgets with these in mind. Power The Irtysh River is one of the major river systems of the world, with seasonal water flow, subject to snow melt. Kazakhstan alone has 3 hydroelectric power plants along its length with the Bukhtarma plant, leased and operated by Kazzinc, being one of them. Its annual generation capacity is 2.4 billion kWh, and Bukhtarma’s power is used for balancing the national grid [13]. Average power costs after distribution are about USD$0.03 per kWh [13]. Transport Geographically UKMC is farther from the sea than just about any place on earth, which is a significant disadvantage for both custom concentrate purchases and metal sales to consumers. The nearest sea ports are either in Russia on the Black Sea coast, or in China on the Pacific Ocean’s west coast. Many thousands of kilometres of rail travel are required in either direction. This factor has driven some of the development initiatives at UKMC. To minimise the impact of logistics costs on the business it is most attractive to purchase concentrates with the highest metal value per tonne, and this has encouraged UKMC to look beyond base metals and become an integrated precious metal producer.

Conclusions The Ust-Kamenogorsk Metallurgical Complex (UKMC) has a long and proud history in sulphide smelting for base metals and precious metals. The foundations have been put in place to achieve the environmental compliance and the workplace productivity improvements that a prosperous future will demand. There has been dynamic growth at UKMC over the past twenty years and the current team is proud of the company’s achievements in this period. An important role will exist for responsible sulphide smelting for a long time to come, and UKMC wants to be part of that future. Acknowledgements The authors would like to acknowledge Kazzinc Ltd for approving data and images for publication. The authors would like to thank Kazzinc Ltd for permission to publish this paper.

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References 1. Nikolić B, Nikolić S, Vujačić V, Trajković S (2014) Development of the production of lead and precious metals in central Asia. In: Undeground Mining Engineering—Podzemni Radovi, vol 24. University of Belgrade, pp 73–79 2. Kennedy KH (1986) Mining Tsar: The Life and Times of Leslie Urquhart. Allen & Unwin 3. Blainey G (1960) Mines in the Spinifex—The Story of Mount Isa Mines. Angus & Robertson 4. Uatkhanov Y (2016) Kazzinc: Human Capital as Key Competitive Advantage. Astana Times 5. Kulbayeva A (2016) Mining and Metals Industry of the Republic of Kazakhstan 2015: Analysis of the Key Economic Indicators. Rating Agency of the RFCA. www.rfcaratings.kz 6. Chadwick J (2009) Kazakhstan: a ‘Market Economy’. Int Min, 8–14 7. Avrachov A, Saltykov P, Strecker JJF, Dobersek A, Azekenov TA, Moskalenko SN (2015) Design & construction of a copper plant based on modern & environmental-friendly processes. In: Proceedings of EMC 2015, Dusseldorf, pp 1–12, GDMB 8. Burrows AS, Alvear GRF, Tynybayev AT (2013) Smelting of Kazakhstan Concentrates at Ust–Kamenogorsk Using a Copper ISASMELT™ Furnace. In: Proceedings of Copper 2013, Santiago, Chile, vol. III, pp 39–48, IIMCH 9. Avrachov A, Saltykov P, Strecker JJF, Dobersek A, Azekenov TA, Moskalenko SN (2015) ISASMELT™-based reconstruction of a lead plant. In: Proceedings of EMC 2015, Dusseldorf, pp 267–272, GDMB 10. Burrows AS, Azekenov TA, Zatayev R (2015) Lead ISASMELT™ operations at Ust– Kamenogorsk. In: Proceedings of Pb-Zn 2015, vol 1., pp 245–256, Dusseldorf, GDMB, 2015 11. Nakupov Zh (2004) Refining & manufacturing perspectives in the CIS: the view from Kazakhstan. In: Proceedings of the LBMA Bullion Forum, Moscow, pp 63–68, LBMA 12. 2017 Full Year Results. http://www.nyrstar.com/∼/media/Files/N/Nyrstar/results-reports-andpresentations/english/2018/2017-full-year-results-release.pdf 13. Mills D, Howard B (2014) Electricity regulation in Kazakhstan: overview. Energy and Natural Resources Multi-Jurisdictional Guide 2014, Thomson Reuters

Trace Metal Distributions in Nickel Slag Cleaning Niko Hellstén, Pekka Taskinen, Hannu Johto and Ari Jokilaakso

Abstract To capture efficiently the valuable trace metals from nickel slag using electric furnace, it is important to study their distributions between the slag and metallic phases in the furnace. It is impossible to calculate accurately these distributions without experimental measurements. Therefore, in this work, selected trace metal distributions and phase equilibria between K2O containing iron-silicate slags and a metallic Ni-Fe-Cu alloy in nickel slag cleaning furnace conditions were studied. The experimental method developed and applied during the work, involved a modified quenching technique that included equilibration of the samples in semi-sealed quartz ampoules in an inert atmosphere and metallic Fe saturation. The use of the semi-sealed quartz ampoule prevents the escape of volatile elements from the sample during equilibration. Chemical compositions of the phases and the trace elements were analysed by EPMA. From the measured compositions, the trace metal distribution coefficients between the molten Ni-Fe-Cu alloy and slag were calculated. Keywords Ni slag cleaning Semi-closed ampoule



Distribution of ag



Au



Sb and pb

Introduction Most pure nickel is produced from sulphidic ores by smelting. Typically, in the smelting of Ni-bearing concentrate, economically important amounts of Ni as well as metal values (for example Co, PGMs, Ag and Au) remain in the slag phase as entrained matte (or as dissolved oxides). To recover the metal values, the slag is cleaned in an electric furnace under reducing conditions [1]. The aim in this process N. Hellstén (✉) ⋅ P. Taskinen ⋅ A. Jokilaakso School of Chemical Engineering, Aalto University, P.O. Box 16200, Espoo, Finland e-mail: [email protected]fi H. Johto Outotec Research (Pori), Kuparitie 10, 28101 Pori, Finland © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_30

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is to reduce the oxidized and dissolved metals so that they would report to a liquid metallic phase consisting mainly of Fe and Ni. Optimization of the metal recovery requires experimental determination of the distribution of the metals between slag and metallic phase at the process conditions. However, due to the high vapour pressures of some of the metal values, for example Ag, Pb and Sb, they may volatilise under the experimental conditions [2]. Therefore, experimental determination of their distributions in the molten phases is challenging. Development of an experimental method to study the behavior of the volatile elements would be of value in the research. The use of sealed ampoules in experiments is an established method. However, the downside of a sealed ampoule is that control of the gas atmosphere is difficult. Recently, Tirronen et al. [3] conducted experiments using an ampoule design with a hole, to enable gas atmosphere from outside the ampoule to equilibrate with the atmosphere inside. The aim of this work was to develop and test a semi-closed ampoule technique to study equilibrium behavior of the volatile elements Ag, Pb and Sb. We made a hypothesis, that the semi-closed ampoule forms a kinetic barrier for the saturated gas phase above the liquid slag and alloy so that the volatile elements remain in the condensed phases during equilibration. In addition, the saturated gas phase should prevent further volatilization. An advantage of this method is that the gas atmosphere inside the ampoule can be controlled externally by, e.g., CO-CO2 mixtures. Using a semi-closed ampoule, the distributions of volatile trace elements Ag, Pb and Sb and non-volatile Au between K2O containing iron-silicate slag of fixed Fe/ SiO2 ratio and a liquid Cu-Fe-Ni metal alloy, at solid Fe alloy saturation, were studied at 1400 °C by equilibration in reducing conditions and quenching. The previous research on the behavior of many impurity elements available in the current experimental conditions is scarce [4] and it is evident that the published data are biased by the trace element volatilization [5]. The distributions of these elements between the slag, a liquid Cu-Fe-Ni alloy and solid Fe alloy were studied as a function of the Fe/Ni ratio in the liquid alloy. The used experimental conditions were chosen because they closely resemble the industrial process [6]. Because Ag, Pb and Sb have been found to volatilise rapidly under the experimental conditions [7], a novel semi-closed ampoule technique was developed and tested in this work.

Materials and Methods The Cu-Fe-Ni alloy, iron silicate slag and Fe foil crucibles used in the experiments were prepared from high-purity chemicals (Table 1). The semi-closed ampoule was made of fused pure quartz.

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Table 1 Purities and suppliers of the chemicals used Chemical

Purity (wt%)

Supplier

Cu Ni Fe Fe foil, 0.25 mm (for crucibles) SiO2 Fused quartz glass (for ampoules) Fe2O3 K2CO3 (as a source of K2O) Ag Au Pb Sb

99.999 99.996 99.99 99.99 99.99 99.99 99.99 99.5–100.5 99.99 99.96 99.999 99.999

Alfa-Aesar Alfa-Aesar Alfa-Aesar Alfa-Aesar Umicore Finnish Special Glass Alfa-Aesar Sigma-Aldrich Alfa-Aesar Alfa-Aesar Alfa-Aesar Cerac

Synthesis of Cu-Fe-Ni Alloy We synthesized four Cu-Fe-Ni alloys of varying Fe/Ni ratio for the equilibration experiments, see Table 2. According to the calculated phase diagram [8], iron and nickel form a single solid alloy at 1400 °C that covers the whole concentration range of the system at low copper concentrations. To obtain a molten metallic phase in the experiments, Cu was added to the alloy. In contrast, in slag cleaning in real industrial process conditions the metallic phase, forming electric furnace matte, is molten due to alloying elements lowering its melting temperature [9]. For synthesis, we weighed and mixed pure metal powders of Cu, Fe and Ni in the desired ratios. Subsequently, the mixtures were heat treated in a horizontal tube furnace at 1400 °C in pure Al2O3 crucibles supported by an alumina boat. We used an atmosphere of 99% Ar (99.99%, Linde Aga) and 1% of H2 (99.99%, Linde Aga) in the synthesis to avoid oxidation of the metals. The furnace was heated to 1400 °C at a rate of 4 °C/min, held at the target temperature overnight and cooled to room temperature 4 °C/min. From a visual observation, the alloys containing 40 wt% and 30 wt% of Ni looked more completely molten than the ones containing 60 wt% and 50 wt% of Ni, Table 2 Normalised SiO2, FeO and K2O concentrations in the slag in equilibrium with the different Cu-Fe-Ni alloy compositions, calculated from the EPMA analysed Si and K concentrations Alloy

SiO2 (wt%)

StDev

FeO

StDev

K2O (wt%)

StDev

I (40/60) II (50/50) III (60/40) IV (70/30)

30.86 36.64 34.18 33.74

3.62 4.10 1.00 1.05

67,80 61,85 64,09 64,89

3,50 5,17 1,21 1,12

1.34 1.51 1.74 1.37

0.47 0.61 0.22 0.40

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respectively. The samples containing more nickel contained edged crystal shapes, which were probably solid precipitates formed in the slow cooling process. Table 3 shows the average elemental compositions of the synthesized Cu-Fe-Ni alloys after melting, as measured by EDS.

Preparation of Sample, Fe Crucible and the Semi-closed Ampoule An iron crucible (length 30 mm, outer diameter of 5.8 mm or less) was prepared from 0.25 mm thick pure Fe foil by rolling and wrapping into a crucible shape. The bottom of the ‘crucible’ was pressed tight using pliers. The iron saturated iron-silicate based slag mixture was prepared by mixing Fe, Fe2O3, SiO2 and K2O (obtained from a pre-reacted and premelted SiO2-K2CO3 mixture) powders to form a slag, which contains 32 wt% of SiO2 and 2 wt% of K2O in equilibrium. In addition, 1 wt% of Ag, Au, Pb and Sb were added as powders into the slag mixture. The Cu-Fe-Ni alloy and slag were inserted into the Fe crucible. Because the alloy and slag react strongly with the foil at the experimental temperature, pure Fe flakes were added in the bottom of the crucible to protect its integrity during experiments. Pure Fe, the Cu-Fe-Ni alloy and slag containing the trace elements each weighed 0.2 g, and thus their corresponding weight ratios were 1:1:1. The semi-sealed ampoule was formed from a quartz tube (8 mm OD, 6 mm ID; Heraeus: HSQ 300). A 1.9 mm diameter hole was drilled into the ampoule 40 mm above the bottom of the tube using a dentist’s drill (NSK Presto Aqua II, NSK, Japan) and a natural diamond ball point drill head (Intensiv, Switzerland). The iron foil crucible containing the sample was inserted into the tube. Subsequently, the quartz tube was cut to 50–80 mm length and a hook for suspension was formed on the top-end using a hydrogen-oxygen torch. This operation left the drilled hole as the only opening into the ampoule to the surrounding furnace atmosphere. Figure 1 shows a schematic of the semi-sealed ampoule containing the Fe crucible and the sample inside. In this method, we assumed that the semi-sealed ampoule forms a kinetic barrier for the volatile elements. During experimentation, the gas phase above the liquid slag saturates with the volatile elements. The small hole decreases the removal rate of the saturated gas phase, by Knudsen-like gas diffusion, while the saturated gas phase reduces transfer of the volatile elements from the condensed phases into the gas flow. Thus, it was expected that the volatile elements equilibrate between slag and alloy, and remain in the sample after quenching in sufficient amounts for in situ analysis with EPMA (Electron probe X-ray microanalysis) or LA-ICP-MS (Laser ablation-inductively coupled plasma-mass spectrometry). In equilibrium, the ratio and thus the distribution coefficient of the metals between the Cu-Fe-Ni alloy, slag and metallic Fe, remains the same regardless of evaporation of the metals if the diffusional mass transfer in the slag and alloy are faster than the volatilisation. Figure 2 illustrates the new concept.

Fe/Ni original

43/57 51/49 60/40 70/30

Alloy # target Fe/Ni

I (40/60) II (50/50) III (60/40) IV (70/30)

65/35 71/29 74/26 77/23

Fe/Ni equilibrated Initial 40.6 35.3 30.8 20.0

Stdev 0.39 0.39 1.26 2.01

Initial 30.9 37.5 45.9 46.7

Equilibrated 55.4 58.2 59.1 58.7

Ni (wt%)

Fe (wt%) Equilibrated 29.5 23.4 20.8 17.8

Table 3 Targets and the measured Fe/Ni ratios (w/w) and compositions of the synthesized Cu-Fe-Ni alloys

Stdev 0.30 0.22 1.44 2.66

Initial 28.6 27.2 23.2 33.3

Cu (wt%) Equilibrated 15.1 18.4 20.1 23.4

Stdev 0.55 0.48 2.13 1.09

Trace Metal Distributions in Nickel Slag Cleaning 383

384 Fig. 1 Schematic of the cross-section of the semi-sealed ampoule, iron foil crucible and the slag-alloy sample

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01.9 mm hole

Slag

Quartz ampoule

Cu-Fe-Ni alloy

Fe foil

When pMe,sl, pM,M < pMe,inside, volatilization of the metal/oxide to the surrounding furnace atmosphere will occur but it is assumed to be slow compared to vaporization processes on the slag surface. When pMe,inside < pMe,outside, the metal/ oxide containing gas moves outside the ampoule through the hole. In this experimental arrangement, the molten alloy-solid alloy equilibrium with FeO in the slag forms the prevailing oxygen pressure in the ampoule.

Experimental Method The equilibration experiments were conducted in a vertical tube furnace at 1400 °C under N2 (Nitrogen 5.0, ≥ 99.999%, Aga Linde) atmosphere using a similar setup and method as described earlier [10]. The semi-sealed ampoule containing the sample was introduced into the furnace work tube from its lower end and suspended there with a 0.5 mm Pt-wire. The ampoule was kept in the low end while the furnace was sealed, and the gas atmosphere flushed with flowing N2 for 30 min, at 200 mL/min flow rate. After flushing, the ampoule was raised to the hot zone of the furnace, equilibrated for 4 h and quenched into a 0 °C ice-water mixture. During experiments, temperature of the sample was monitored using a calibrated S-type Pt-Pt90/Rh10 thermocouple. Fig. 2 Schematic of the gas phase arrangement and the expected behavior in the experimental technique used

Outside:

Inside the ampoule:

unsaturated

saturated

pMe

pMe inside Slag

pMe,sl Liquid

pMe,

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Time required for equilibration was determined experimentally as 4 h based on changes in the slag composition. After quenching, the sample was dried, mounted into epoxy and polished using wet metallographic methods. The sample was made conductive carbon coating by evaporation. Microstructure observations and preliminary analysis of chemical compositions of the phases were conducted using SEM-EDS (LEO SEM 1450—Oxford Instruments X-Max 50 mm2 EDS + INCA). The accurate chemical phase compositions were carried out by EPMA (Cameca SX100). The EPMA was operated using 20 kV accelerating voltage, 40 nA beam current and 50 µm–100 µm beam diameters. A PAP matrix correction program supplied by the manufacturer was used to correct the analytical results. The external standard materials and analysed lines used in the EPMA analysis were Fe Kα and O Kα (hematite), Si Kα (quartz), K Kα (sanidine), Cu Kα (Cu), Ni Kα (Ni), Ag Lβ (Ag), Sb Lα (Sb-telluride), Au Lα (Au) and Pb Lα (galena).

Results and Discussion Initially, a series of test experiments were conducted to find working experimental parameters. Time required for the attainment of equilibrium were determined by a time series. Criteria for a successful experiment was that the three phases (slag, liquid Cu-Fe-Ni alloy and solid Fe-rich alloy) were present in areas large enough for microanalysis in situ. For possible laser ablation studies, the optimal results would require 100 µm × 100 µm sized areas in minimum. Finally, we analyzed with EPMA the chemical compositions of the liquid Cu-Fe-Ni alloy, molten slag and metallic, solid Fe-rich phases in eight quenched samples—two from each Cu-Fe-Ni alloy of different Fe:Ni ratios. Eight individual measurements from each phase but different locations were taken from well-quenched areas in each sample. The total values in all the original measurements were between 97 wt%–101 wt%, which are considered good. In each sample, the concentrations of Ag, Sb and Pb of about 1 wt% could be analysed from the molten Cu-Fe-Ni alloy and much smaller concentrations in the slag phase. Composition of the gas phase was not analysed in this study as the focus was on the distribution of trace elements between the condensed phases.

Trace Element Distributions From the measured concentrations of Ag, Au, Sb and Pb by EPMA the distribution coefficients between the liquid Cu-Fe-Ni alloy and molten iron silicate slag were calculated as LMe/S(M) = (wt%M)/[wt%M]. Figure 3 shows that all of the studied elements report strongly to the molten Cu-Fe-Ni alloy phase in nickel slag cleaning conditions, near the saturation of the alloy with pure iron. This is in agreement with the initial hypothesis.

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Fig. 3 (a)–(d) Experimental distribution coefficients of Ag, Au, Pb and Sb between the liquid Cu-Fe-Ni alloy and iron silicate slag as a function of Fe concentration of the alloy at 1400 °C

In addition, Ag shows a slight increasing trend with increasing iron content in the alloy, while in the case of Pb a similar increase is more pronounced. The case for Sb shows a slight decrease as the iron content in the alloy increases. Surprisingly, the results for Au show large scatter. One explanation for this may be that the detected concentrations of Au, and the other traces in the slag, were clearly below the detection limits of EPMA (Ag: 400, Au: 610, Pb: 320, and Sb: 190 ppm), causing severe uncertainty in the analysis. Due to the large values of the distribution coefficients, the EPMA measured values for slag for example, indicating very low concentrations in the slag, may be rather imprecise and indicate a lower value of the true distribution coefficients. An improvement will be attained with more sensitive analytical techniques, i.e. LA-ICP-MS [7] to be used for further studies of the present samples in a near future. Table 2 shows the normalized SiO2 and K2O concentrations of the slag phase analysed from the samples. The values obtained for each slag in Table 2 are averages from two different samples and eight different analysis locations in each sample. Table 2 shows that the SiO2 and K2O concentration are close to the estimated target values of 32 wt% of SiO2 and 2 wt% of K2O, respectively. Deviations between the target and measured SiO2 concentrations may be caused by uncertainties from the preparation of the original sample mixtures. For example, SiO2 was added as pure SiO2 and as a mixture containing approximately 90% SiO2 and 10% K2O. Thus, control of the exact amounts of SiO2 and K2O in the mixture was unobtainable. In addition, the concentration of K2O in the slag affects the solubility of SiO2. Thermodynamically, as the system has two degrees of freedom at the experimental conditions, it should be possible to fix the SiO2 concentration of the slag. However, as the activity of Fe and FeO is fixed, it may affect the equilibrium

Trace Metal Distributions in Nickel Slag Cleaning Fig. 4 Quenched microstructure of a sample containing the molten Cu-Fe-Ni alloy III (Table 3)

387

Fe foil Cu-Fe-Ni alloy

Slag

concentration in each case. In this work, the accuracy of the EPMA measurements was determined by measuring a standard sample (Sect. 2.3). Acceptable deviation from the nominal value of the standard sample was 1 wt%. Due to the high totals of the original measurements and the standards used, the accuracy of the EPMA measurements in this work is considered good. During experimentation, it was discovered that the thin iron foil crucible reacted with the molten alloy sample forming holes and causing the slag to react locally with the quartz ampoule. Thus, to ensure that the sample was in solid Fe alloy saturation, criteria for a successful experiment was that solid Fe-rich alloy coexisted with the molten slag and Cu-Fe-Ni phases after experiment. The Fe-rich alloy contained 0.1–9.2 wt% Cu and 0.1 to 21.6 wt% Ni. In addition to this phenomenon, quenching of the molten Cu-Fe-Ni alloy into a single homogeneous phase was found to be very challenging. The solidifying Cu-Fe-Ni alloy often segregated into a Cu-rich phase and Fe-Ni rich phase in quenching, resembling the presence of miscibility gap in the copper-iron binary system [11] and its extension to the ternary alloys. It is possible that the presence of trace elements, such as lead and possibly silver, enhances this behavior [11]. Moreover, bright spots in the Cu-rich areas containing larger concentrations of the trace elements than the iron-rich matrix (Fig. 4).

Liquid Cu-Fe-Ni Alloy Phase Visual differentiation of the Cu-Fe-Ni molten alloy from the solid metallic iron alloy after quenching in BSE image was sometimes difficult because the metals start to form an alloy, generating the tie-line, between the solid and liquid alloys. At the same time, the composition of the liquid alloy moves towards the Fe corner of the ternary system. In addition, high-purity iron foil dissolves nickel and copper. According to phase equilibria of the Cu-Fe-Ni system, pure Fe is unstable at 1400 °C [11]. Instead, it dissolves into a liquid Cu-Fe-Ni alloy and a (γFe, Ni) solid

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phase which constitute the equilibrium phase-assemblage. The (γFe, Ni) phase is an extended solid solution able to dissolve Cu. Initially, the solid Fe and the (γFe, Ni) alloy form a diffusion couple and the equilibrium is first obtained in a reaction layer next to their interphase. Due to this phenomenon, a profile from the boundary between the molten Cu-Fe-Ni alloy and slag towards the pure Fe-foil was measured to determine the size of the molten Cu-Fe-Ni phase and the extent of diffusion and formation of the (γFe, Ni) solid solution phase. Average oxygen amounts of 0.27, 0.55, 0.57 and 0.52 wt% were analysed by EPMA in the Cu-Fe-Ni alloys I-IV, respectively, as traces of the sample preparation and forming a typical surface contamination by atmospheric interactions.

Conclusions A novel method to study the behavior of easily volatile elements in nickel slag cleaning electric furnace conditions close to iron saturation was developed. The results indicate that the semi-closed ampoule may be used to study these volatile elements, as significant fractions of the initial, weighed amounts of the volatile Ag, Pb and Sb could be found in the equilibrated slag-alloy samples. This result indicates that the semi-closed ampoule method adopted is very useful in studying the behavior of these easily volatilizing trace elements in various slag-metal and slag-matte systems. Because the method is based on the formation of a kinetic barrier to slow down the movement of the saturated gas phase from the ampoule and thus decreasing the rate of volatilization, the concentrations of these elements in the gas phase are irrelevant in the calculation of the distribution coefficients between the condensed phases. In equilibrium, these coefficients are independent of the amounts of the materials present in the phases. Distribution of Ag, Au, Pb and Sb between liquid Cu-Fe-Ni alloy and molten iron silicate slag were studied at 1400 °C in metallic Fe saturation. The experimental series were carried out in inert atmosphere and nickel concentrations in the molten alloy ranged from 20 to 40 wt%. It was found that all the studied elements distribute preferentially to the liquid metallic Cu-Fe-Ni alloy phase under the studied conditions. Quenching the liquid Cu-Fe-Ni alloy caused uncertainties in the study of this particular system using a quartz ampoule where the cooling rates were much lower than with low-mass, open samples. The obtained distribution coefficients for volatile elements, e.g. antimony, between metal and slag are lower than reported earlier [5] by about one order of magnitude. This may be a result of systematic depleting the slag phase from the trace element due to volatilization, which is very difficult to discover in the crucible equilibration-sampling measurements. Acknowledgements The authors would like to thank Boliden Harjavalta and CMEco project of Tekes for the financial support of this project. Mr. Lassi Pakkanen and the Finnish Geological Survey are acknowledged for the EPMA measurements.

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References 1. Crundwell FK, Moats MS, Ramachandran V, Robinson TG, Davenport WG (eds) (2011) Extractive metallurgy of nickel, cobalt and platinum group metals. Appendix E—recovering nickel-, copper-, cobalt- and platinum-group elements from slag, Elsevier, Oxford, pp 567–574 2. Yazawa A (1980) Distribution of various elements between copper, matte and slag. Erzmetall 33:377–382 3. Tirronen T, Sukhomlinov D, O’Brien H, Taskinen P, Lundström M (2017) Distributions of lithium-ion and nickel-metal hydride battery elements in copper converting. J Clean Prod 168:399–409 4. Chen J, Allen C, Hayes PC, Jak E (2016) Experimental study of Slag/Matte/Metal/Tridymite four phase equilibria and minor element distribution in “Cu-Fe-Si-O” system by quantitative microanalysis techniques. In: Advances in molten slags, fluxes, and salts: Proceedings of the 10th international conference on molten slags, fluxes, and salts (Ed. Reddy R., Chaubal P., Pistorius P.C., Pal U.). TMS, Warrendale (PA), pp 961–970 5. Pagador RU, Hino M, Itagaki K (1999) Distribution of minor elements between MgO saturated FeOx-MgO-SiO2 or FeOx-CaO-MgO-SiO2 slag and nickel alloy. Mater Trans, JIM 40(3):225–232 6. Taskinen P, Seppälä K, Laulumaa J, Poijärvi J (2001) Oxygen pressure in Outokumpu flash smelting furnace. Part II. The DON Process. Trans IMM, Sect C, 110(2):101–108 7. Avarmaa K, O’Brien H, Johto H, Taskinen P (2015) Equilibrium distribution of precious metals between slag and copper matte at 1250–1350 C. J Sustain Metall 1(3):216–228 8. Davies RH, Dinsdale AT, Gisby JA, Robinson J, Martin SM (2002) MTDATA— thermodynamics and phase equilibrium software from the national physical laboratory. Calphad 26:229–271 9. Mäkinen T, Taskinen P (2008) State of the Art in Nickel Smelting: Direct Outokumpu Nickel Technology. Trans IMM Sect C 117(2):86–94 10. Hellstén N, Hamuyuni J, Taskinen P (2015) High-temperature solubilities in the Cu-O-MgO system at metallic copper saturation. In: Proceedings European Metall. Conference EMC 2015, Vol. 1. GDBM, Clausthal-Zellerfeld, pp 47–60 11. Dreval LA, Turchanin MA, Agraval PG (2014) Thermodynamic assessment of the Cu-Fe-Ni system. J. All. Comp. 587(1):533–543

Case Study on the Application of Research to Operations—Calcium Ferrite Slags Stanko Nikolic, Denis Shishin, Peter C. Hayes and Evgueni Jak

Abstract The Top Submerged Lance (TSL) technology, developed in the 1970s, is now widely used for the processing of a range of materials. TSL technology for continuous converting was first patented in the 1990s. The process is based on the use of calcium ferrite slag. Although this slag system had been applied elsewhere the phase equilibria had not been thoroughly investigated. This lead to a collaboration between the Process Technology group of Mount Isa Mines, now part of Glencore, and the Pyrometallurgy Innovation Centre (PYROSEARCH) at The University of Queensland. Through multiple research programs this complex system was successfully investigated. The results were then implemented within thermodynamic modelling tools. This combined new knowledge was then applied to the design and industrial implementation of the TSL continuous converting technology, ISACONVERT™. This paper describes the key findings of the research and how this was applied to the industrial implementation of the technology. Keywords Slags Copper converting



Calcium ferrite



Lime ferrite



Liquidus

Introduction Development of the top submerged lance (TSL) technology began in the 1970s and started with investigations on how to improve tin smelting processes. This work lead to the development of the SiroSmelt lance, which through collaboration S. Nikolic (✉) Glencore Technology Pty. Ltd., Level 10, 160 Ann Street, Brisbane, QLD 4000, Australia e-mail: [email protected] D. Shishin ⋅ P. C. Hayes ⋅ E. Jak School of Chemical Engineering, PYROSEARCH—Pyrometallurgy Innovation Centre, The University of Queensland, Brisbane, QLD 4072, Australia e-mail: [email protected] © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_31

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between the Commonwealth Scientific Industrial Research Organisation (CSIRO) and Mount Isa Mines, now part of Glencore, by the 1990s had expanded the technology to industrial implementations in copper, nickel and lead smelting [1]. During this period the application of TSL technology to the continuous converting of copper matte with lime/calcium ferrite slag was first patented [2]. This patent was based on an extensive pilot plant program with the technical basis provided by earlier phase equilibria work completed by Yazawa [3, 4] and by Takeda [5] in the 1980s. However, understanding of the fundamental compositional and physical properties of the molten slag, such as phase equilibria and viscosity, was missing from the earlier research and would be vital for the process implementation, optimization and control of this TSL technology. This lead to a collaborative effort beginning in the late 1990s between the Process Technology group of Mount Isa Mines, now part of Glencore, and the Pyrometallurgy Research Centre (PYROSEARCH) at The University of Queensland for the purpose of investigating the fundamental phase equilibria of the lime/calcium ferrite slag system, and in particular the “Cu2O”– FeO–Fe2O3–CaO system in equilibrium with liquid copper metal.

Methodology The initial work on this system at PYROSEARCH focused on applying the rapid quenching electron probe x-ray microanalysis (EPMA) technique previously established by Jak et al. [6–8]. The advantage of the rapid quenching technique is that it allows for the phase assemblages present at equilibration to be retained at room temperature. However, direct application of this technique proved challenging, due to the low viscosity of the lime/calcium ferrite slag, with the majority of the initial experimentation programs focusing on achieving a well quenched slags amenable to EPMA analysis. Using this micro-analytical technique the metal concentrations in the phases could be measured to within ±1wt%. The oxidation state of the iron could not be measured using this technique. Two techniques were then investigated for producing well-quenched slags: using levitated copper droplets and using substrates made from the primary phase [8]. Both these techniques allowed experiments to be completed at metallic copper saturation and eliminated the uncertainties associated with contamination from crucible materials. However, these experiments were undertaken in a neutral gas atmosphere and although they provided useful phase equilibria information they were not able to provide information on the effect of oxygen partial pressure on the system. Knowledge of the effect of oxygen partial pressure on the phase equilibria was required to relate the research to the industrial process control parameters of the sulfur content of the produced blister copper and copper content within the slag. When evaluating the suitability of experimental designs at fixed oxygen partial pressure the primary phase substrate technique was selected as the superior method. An explanation of the vertical tube furnace and glass capillary gas flow meter setup

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Fig. 1 Primary phase substrates used in the present study—a Spinel (Fe3O4), b Dicalcium ferrite (Ca2Fe2O5) and c Cuprite (Cu2O) Nikolic et al. [11]

used for the fixed oxygen partial pressure experiments for this system through the use of using flowing using CO/CO2 or H2/CO2 gas mixtures is contained in the preceding papers Nikolic et al. [9–11].

Substrate Design There were three primary phase substrates used: magnetite/spinel (Fig. 1a), dicalcium ferrite (Fig. 1b) and cuprite (Fig. 1c)—Nikolic et al. [11]. These three substrate shapes maximized the available surface area used for suspending slag whilst minimizing the mass of the substrate and thereby maximizing the cooling rate of the slag on quenching. A key feature common to all of these designs is that the slag is on the outside of the support, held in a thin (50–200 μm) film by surface tension onto the underlying solid substrate.

Results and Discussion Microstructures in the “Cu2O”–FeO–Fe2O3–CaO System in Equilibrium with Liquid Copper Experiments within the “Cu2O”–FeO–Fe2O3–CaO system in equilibrium with liquid metallic copper were conducted at temperatures between 1100 and 1250 °C at 50 °C steps, in a flow of ultra high purity argon gas, and at 1200 and 1250 °C at fixed oxygen partial pressures in the range of 10−5.0–10−6.5atm. Scanning Electron Microscope (SEM) micrographs of equilibrated experimental samples with

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examples of the phase assemblages identified in the study are presented in Fig. 2. These micrographs show, in each case, that well-quenched liquid slag (marked as glassy slag in high magnification inserts) and metallic copper in equilibrium with the relevant solid phases are obtained. Compositional analysis of the liquid phases, with EPMA, was undertaken on the well-quenched glassy slags, shown in the inserts in Fig. 2a–f, which are found at the liquid/quenching medium interface.

Fig. 2 SEM micrographs of typical well-quenched slag samples showing all phase assemblages identified in the study with solid crystal(s) in equilibrium with liquid slag and metallic copper obtained using the primary phase substrate technique Nikolic et al. [12]

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Using this approach the composition of the liquid slag in equilibrium with the solid substrate can be directly measured, and the liquidus surface at fixed temperature and fixed oxygen partial pressure can be determined experimentally. The iron present in the slags is present in both 2+ and 3+ oxidation states, however as indicated above the concentration of these ionic species cannot be readily determined using the EPMA technique.

Liquidus Temperatures Defined by Cu2O Concentration and Fe/CaO Weight Ratio In industrial practice, selection of operating conditions in copper converting furnaces typically involves fixing the Fe/CaO ratio and targeting a fixed copper oxide concentration in the slag phase at a selected temperature. The Fe/CaO ratio and the copper oxide concentration in the slag can be controlled through effective mass balancing of the system by accurately measuring and controlling oxygen inputs, and feed flow rates and compositions. The temperature of the system is the other key process variable, therefore it is important to consider how the copper oxide concentration and Fe/CaO ratio affect the liquidus of the system as detailed in Figs. 3 and 4. The phase equilibria data that were obtained from the study for the “Cu2O”– FeO–Fe2O3–CaO system, from Nikolic et al. [11, 12], have been represented in Fig. 3 as a function of copper oxide concentration in the slag phase and fluxing Fe/ CaO wt ratio. The data indicate that the liquidus is very sensitive to both Fe/CaO ratio and the concentration of dissolved copper oxide in the slag.

Fig. 3 Liquidus temperatures in the “Cu2O”–FeO–Fe2O3–CaO system in equilibrium with copper metal as a function of Cu2O concentration and Fe/CaO ratio in the slag

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Fig. 4 Liquidus in the “Cu2O”–FeO–Fe2O3–CaO system represented as Fe/CaO wt% ratio at different Cu2O concentrations in liquid

The data given in Fig. 3 can be represented in an alternative format. The minimum liquidus temperatures, at each fixed copper oxide concentration, shown in Fig. 4, are based on extrapolations of the experimental liquidus data and the positions of the univariant lines in the system (refer to Fig. 3). The information in Fig. 4 indicates that the minimum liquidus temperatures of the slags will decrease as the copper oxide concentrations increase from 10 to 40wt% and will continue to decrease until the ternary eutectic of the system is reached. The diagram enables the operating window for the slag composition to be readily identified. The bounding primary phase fields are dicalcium ferrite (2CaO.Fe2O3) and spinel/magnetite (Fe3O4). At 1250 °C, for example, the slags are fully liquid in the range Fe/CaO wt ratio between 1.7 and 3.0 at 20wt% Cu2O. At 10wt% Cu2O the range of fully liquid slag reduces to Fe/CaO wt ratio between 2.0 and 2.8.

Application of the Research to Plant Design and Operation This and more recent experimental data produced by the PYROSEARCH group has been incorporated into optimized thermodynamic databases describing gas/slag/ matte/metal/solids phases as described in Shishin et al. [13]. The database is now being used in conjunction with the FactSage computer platform to predict outcomes in copper smelting and converting processes. The detailed explanation of the application of these calculations to copper smelting and converting processes is described in Jak [14]. In the examples given below, the databases have been used to predict and compare the outcomes of three alternative process combinations (see Fig. 5) with

Case Study on the Application of Research to Operations …

Air + O 2 60 vol% O2

(+ Slag/matte separation)

Concentrate 200 t/h SiO2 flux Air + O 2 60 vol% O2

Gas

Slag blow 1

(+ Slag/matte separation)

Gas ISA Smelt (+ Slag/matte separation)

Gas

Slag blow 2 High-Cu slag

Slag Gas ISA Smelt

Air + O 2 60 vol% O2

Matte (1180 °C) 60 wt% Cu

RHF

Reverts SiO2 flux Cu scrap

Reverts SiO2 flux

Reverts, 10 t/h

Gas ISA Smelt

Converter slag (25 °C), all Matte (25 °C), 70 wt% Cu

Slag

397

RHF

CaCO 3 flux Air + O 2 60 vol% O2

Converter slag (25 °C), all SiO2 and CaCO 3 flux Matte (25 °C), 70 wt% Cu Air + O 2 Slag RHF 60 vol% O2

Gas Copper blow Cu >99%

Gas ISA Convert (Calcium slag)

Cu >99% Gas ISA Convert (Calcium ferrous silicate slag)

Cu >99%

Fig. 5 Process flow diagrams showing case studies considered Case 1 ISASMELT™ + Peirce Smith; Case 2 ISASMELT™ +ISACONVERT™ Calcium ferrite; Case 2 ISASMELT™ + ISACONVERT™ ferrous calcium silicate

ISASMELT™ used for smelting and with alternative technologies used for converting (Peirce Smith and ISACONVERT™). Process Block Diagrams showing the sequence of operations and material flows are given in Fig. 5. Case 1 demonstrates the use of the ISASMELT™ smelting stage with fayalite slag in series with conventional Peirce Smith converters (PSC) having fayalite slags. Case 2 is the use of the ISASMELT™ smelting stage with fayalite slag in series with ISACONVERT™, the latter having calcium ferrite based slag. Case 3 is the ISASMELT™ smelting stage with fayalite slag in series with ISACONVERT™ using ferrous calcium silicate based slag. The smelting slag liquidus (projected on to the Temperature vs Fe/SiO2 pseudo-binary system) and normalized converting slag compositions (projections on to the “FeO”–CaO–SiO2 pseudo-ternary system) are illustrated in Figs. 6 and 7 respectively. This example assumes inputs of 200 t h−1 of 25.0%Cu concentrate, silica and/or lime flux, air and tonnage oxygen (making up a blast of 60vol.% O2) in the smelting stage to produce 60%Cu matte for the Peirce Smith converters and 70%Cu matte for the ISACONVERT™ cases. The ISASMELT™ is operated at 1180 °C in all cases. For case 1 the matte is charged in molten state to the Peirce Smith converters, as is the conventional practice. For cases 2 and 3 the matte is cooled to 25 °C before feeding to the ISACONVERT™ furnaces operated at 1250 °C. Details of the feed compositions are provided in Table 1, and a summary of inputs and calculated outcomes from the thermodynamic modelling of the case studies is given in Table 2. Most of the predicted input and output values are listed in Table 2 and are self-explanatory. The concentration of Cu chemically dissolved in slag (row 18) is predicted with the thermodynamic model, and the concentration of Cu physically entrained in slag (row 19) was assumed to be 0.3–0.4wt% for

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Fig. 6 Predicted liquidus compositions for smelting slags for cases 1–3

Fig. 7 Predicted (normalised) converter slag compositions for cases 1–3, projected onto the “FeO”– CaO–SiO2–Cu–O System at 1200 °C and PO2 = 10−5 atm

analysis purpose. The Cu recovery from total input for each of the cases are given in row 22. Row 21 marked “Heat losses [MW]” refers to the net heat generated by the reactor that is dissipated to the environment, through the furnace linings or available for heating supplementary feed; the calculations do not take into account the details of the different reactor designs. The calculations do indicate however that the enthalpy available for the ISASMELT™/ISACONVERT™ combination using the ferrous calcium silicate converter slag is significantly lower than the

Concentrate wt% SiO2 flux wt% Lime flux wt%

0.20 0.00 0.00

Pb

25.00 0.00 0.00

Cu

28.00 1.40 0.70

Fe

30.00 0.00 0.50

S 5.00 92.00 3.00

SiO2

Table 1 Solid feed compositions used for case studies

0.90 0.50 1.00

Al2O3 0.80 1.68 47.60

CaO 1.00 0.10 0.00

MgO 0.100 0.000 0.000

Zn 0.110 0.000 0.000

Ni 0.010 0.000 0.000

Sn 0.150 0.000 0.000

As

0.003 0.000 0.000

Sb

0.012 0.000 0.000

Bi

50 0 0

Ag (ppm)

5 0 0

Au (ppm)

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Table 2 Summary of calculated outcomes of smelter and converter case studies ISASMELT™ −PSC

ISASMELT™− ISACONVERT™

ISASMELT™ −ISACONVERT™

Smelt

Smelt

Smelt

PSC

Convert CaFe slag

Convert CaFeSi slag

1

2

3

4

5

6

7

8

1

Concentrate or Matte feed [t/h]

200

81

200

69

200

69

2

Cu in Concentrate or Matte feed [wt%]

25

60

25

70

25

70

3

SiO2 flux added to the concentrate [t/h]

22.5

9.7

40.2

0.0

35.6

9.6

4

CaCO3 flux [t/h]

0.0

0.0

0.0

4.5

0.0

7.2

5

Recycling slag added to the feed [t/h]

10.0

6.5

12.6

0.0

26.2

0.0

6

Reverts added to the feed [t/h]

0.0

13.0

0.0

0.0

0.0

0.0

7

Fe/SiO2 in the slag [wt%]

1.40

1.45

1.20

36.4

1.2

0.00

8

Fe/CaO in the slag [wt%]

22.4

79.3

12.7

2.3

10.0

0.00

9

Mechanical (physical) dust carry over %

1.00

n/a

1.00

0.50

1.00

0.50

10

Dust added to the fresh feed [t/h]

0.00

0.00

0.00

0.00

0.00

0.00

11

Total feed [t/h]

282

251

293

183

299

156

12

O2 efficiency [%]

100.

100

100

100

100

100

13

O2 enrichment [vol.%]

60.0

24.7

60.0

60.0

60.0

60.0

14

Total blast [Nm3/h]

59,208

82,491

66,381

21,697

64,977

23,310

15

Total output matte or blister [t/h]

81

69

49

69

48

16

%Cu in output matte or blister

60.00

70.00

98.95

70.00

99.15

17

Total slag output [t/h]

97

32

130

13

133

27

18

% Cu chemically dissolved in slag

0.42

4.1

0.44

20.00

0.42

19.99

19

Cu physically entrained in slag (assumed) [wt%]

0.30

0.35

0.39

0.35

0.39

20

T [oC]

1180

1200– 1250

1180

1250

1180

1250

21

Heat losses [MW]

−3.28

?

−8.34

−8.74

−3.34

−3.71

22

Cu recovery from total input [wt%]

0.97

?

0.97

0.99

0.97

0.99

23

Availability of furnace

0.90

0.90

0.90

0.90

0.90

0.90

99.54

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ISASMELT™/ISACONVERT™ combination with calcium ferrite slag. This appears to be associated with the higher slag mass generated in the ferrous calcium silicate converter and the higher flux additions required in this case. On the other hand the heat losses associated with the ISACONVERT™ calcium ferrite based reactor will be greater due to the need to use freeze lining technologies to maintain the integrity of the furnace lining. Although the continuous processes appear to give slightly smaller total production rates this is based on the same availability of all process units and should therefore be viewed with caution. The potential advantages of the continuous processes are that they reduce the proportion of downtime associated with batch processing, delays in material transfers and maintenance and repairs. With the development of accurate thermodynamic databases this approach can be expanded and further used to accurately compare the effect of process variables such as, concentrate feed rate, concentrate grade, impurity content of feed, matte grade, oxygen enrichment on process outputs such as, metal recovery/losses, flux requirements, gas volume and composition, impurity distribution and partitioning all critical to the industrial operating parameters of the process.

Conclusions A top submerged lance technology has been developed for the converting of copper matte to blister copper. The development and implementation of this technology has been supported by fundamental studies on the phase equilibria and, in particular, the liquidus surfaces in the “Cu2O”–FeO–Fe2O3–CaO system in equilibrium with liquid copper in the range of compositions relevant to industrial practice. The project has demonstrated the need for fundamental research that supports industrial research and development. The research has provided accurate information on the chemical behavior of these complex chemical systems, information that was not previously available, greatly assisting understanding the behavior of the processes and enabling the reliable process design and control in industrial practice. Acknowledgements The authors would like to thank the following for their assistance in this project: Ms Jie Yu, Mr Adrian Riding and Dr. Baojun Zhao from the Pyrometallurgy Research Centre (PYROSEARCH) within the University of Queensland; Alistair Burrows, Bill Errington, James Edwards, Martin Bakker and Philip Arthur from Xstrata Technology for useful discussions on this topic; Mr Ron Rasch, Mr Kim Sewell and Ms Ying Yu from the Centre for Microscopy and Microanalysis (CMM) within the University of Queensland. Xstrata Technology and the Australian Research Council provided financial support for this project under the ARC Linkage Program.

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References 1. Burford B (2009) The ISASMELT™ technology package: over 30 years of innovation. AusIMM Bull J Australas Inst Min Metall (1):26–30. February 2009 2. Edwards JS, Jahanshahi S (1999) Copper converting. United States Patent, No. 5,888,270, 30 March 1999 3. Yazawa A, Takeda Y, Waseda Y (1981) Thermodynamic properties and structure of ferrite slags and their process implications. Can Metall Q 20:129–134 4. Yazawa A, Takeda Y (1982) Equilibrium relations between liquid copper and calcium ferrite slag. Trans Jap Inst Met 23:328–333 5. Takeda Y, Nakazawa S, Yazawa A (1980) Thermodynamics of calcium ferrite slags at 1200 and 1300 °C. Can Metall Q 19:297–305 6. Jak E, Hayes PC, Lee HG (1995) Improved methodologies for the determination of high temperature phase equilibria. Korean J Metals Mater 1:1–8 7. Jak E, Hayes PC (2004) Phase equilibria determination in complex slag systems. In: Proceedings of international symposium molten slags and fluxes, capetown. SAIMM, Johannesburg, South Africa, pp 85–104 8. Ilyushechkin A, Hayes PC, Jak E (2004) Liquidus temperatures in calcium ferrite slags in equilibrium with molten copper. Metall Mater Trans B 35B:203–215 9. Nikolic S, Hayes PC, Jak E (2008) Phase equilibria in ferrous calcium silicate slags part i: intermediate oxygen partial pressures in the temperature range 1200–1350 °C. Metal Mater Trans B 39(2):179–188 10. Nikolic S, Hayes PC, Jak E (2008) Phase equilibria in ferrous calcium silicate slags part iii: copper saturated slag at 1250 and 1300 °C at an oxygen partial pressure of 10−6atm. Metal Mater Trans B 39(2):200–209 11. Nikolic S, Hayes PC, Jak E (2009) Experimental techniques for investigating calcium ferrite slags at metallic copper saturation and application to systems “Cu2O”-“Fe2O3” and “Cu2O”CaO at metallic copper saturation. Metal Mater Trans B 40(6):892–899 12. Nikolic S, Hayes PC, Jak E (2008) Liquidus temperatures in the “Cu2O”-FeO-Fe2O3-CaO system at metallic copper saturation. Metal Mater Trans B 40(6):900–909 13. Shishin D, Hayes PC, Jak E (2018) Multicomponent thermodynamic databases for complex non-ferrous pyrometallurgical processes. In: Peter Hayes Symposium Ottawa 14. Jak E (2012) Integrated experimental and thermodynamic modelling research methodology for copper and other metallurgical slags. In: MOLTEN 12, 9th international conference on molten slags, fluxes and salts, Beijing, China, May 2012, paper w77

Kinetics of Oxidation of Pyrrhotite Anastasia Alksnis, Bo Li, Richard Elliott and Mansoor Barati

Abstract Pyrrhotite is an iron deficient, low nickel content sulfide mineral that is commonly associated with pentlandite, a nickel-rich sulfide; and are both main components of the nickel ore mined in Sudbury. Pyrrhotite tailings generated in the concentration of Ni ores are stored in tailing reservoirs and presents serious environmental risks including acid mine drainage. This study investigated the potential of converting pyrrhotite into a valuable resource through the use of fluidized bed roasting, in which the mineral will be oxidized to produce iron oxide and SO2 gas. These products can then be used to create iron for high-grade steel and elemental sulfur. Thermogravimetric analysis of Sudbury pyrrhotite was performed to study the oxidation kinetics, supporting the feasibility of this process. Further, energy and materials balance was carried out to establish the basic parameters of a possible roasting process. Keywords Pyrrhotite



Oxidation kinetics



Fluidized bed roasting

A. Alksnis (✉) ⋅ R. Elliott ⋅ M. Barati Department of Materials Science and Engineering, University of Toronto, Toronto, ON M5S3E4, Canada e-mail: [email protected] R. Elliott e-mail: [email protected] M. Barati e-mail: [email protected] B. Li State Key Laboratory of Complex Nonferrous Metal Resources Clean Utilization, Kunming University of Science and Technology, Kunming 650093, China e-mail: [email protected] © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_32

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Introduction Pyrrhotite (Fe1−xS, x = 0 − 0.125) is a mineral that is largely associated with nickel-copper ores, and is especially abundant in Sudbury, Ontario. Sudbury’s pyrrhotite can contain up to 1 weight percent in nickel dissolved in solid solution in the pyrrhotite matrix. Due to its low nickel content compared to accompanying minerals such as pentlandite ((Fe,Ni)9S), it is typically discarded to increase the smelting yield [1]. The Sudbury area is estimated to have approximately 75–100 million tonnes of pure pyrrhotite accumulated in its tailings ponds, with fresh pyrrhotite being added daily. Despite prior successful attempts to make use of these tailings in the 1950s– 1970s that were forced to shut down due to economic and environmental shifts, there has been no recent initiatives put in place to make use of this material. As a result, its valuable contents are essentially going to waste, while also posing the risk of acid mine drainage to the local environment [1, 2]. In an ongoing study at the University of Toronto, pyrrohite tailings are treated to produce a ferronickel alloy and stoichiometric pyrrhotite known as Troilite, FeS. The ferronickel alloy can be used directly for steel making and FeS can be roasted to produce iron oxide and sulfur dioxide [3]. The products of the roasting can then undergo further refining to produce high purity iron and elemental sulfur or sulfuric acid. Since roasting iron sulfide to iron oxide is a highly exothermic process, the generated heat can be recovered and recycled back into the process to improve the system’s energy efficiency [2]. This study aims to explore how roasting in a fluidized bed can play an integral role in converting pyrrhotite from a waste to resource material. First, mass and energy balance was performed to evaluate the theoretical maximum heat to be recovered, followed by a kinetic study to identify the optimal parameters for rapid oxidation, including gas flow rate, particle size, temperature, and partial oxygen pressure.

Materials and Energy Balance of Roasting Commercial scale processing of pyrrhotite was introduced by Inco’s Iron Ore Recovery Plant in the 1950s and utilized fluidized bed roasters to oxidize the pyrrhotite before further processing to recover its iron and nickel contents. As their operation consisted of similar input and output materials to this study, their roasting specifications can serve as a good basis of what can be expected for processing troilite on a similar scale [4]. Each of Inco’s roasters had a daily capacity of close to 600 tonnes of pyrrhotite concentrate and would be supplied with an air flow of about 1,500 tonnes per day at capacity. The pyrrhotite was fed into the roaster as a slurry with about 25% water by weight. The roasting temperature was held at about 760 °C. The nickel content in the pyrrhotite would oxidize to nickel ferrite, NiFe2O4. Nickel ferrite happens to

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405

have a similar crystal structure to magnetite, Fe3O4, so the roasting parameters were carefully selected to ensure the iron would mainly oxidize to hematite, Fe2O3, as its insolubility for nickel ferrite would simplify the separation downstream [4]. Equations (1) and (2) define the principal reactions used to describe the Inco roasting operation. With inputs consisting of fresh slurry feed and air entering at approximately 15 °C and exiting at 750 °C, along with hot recirculated calcine and cooling water, the theoretical heat input from the roasting operations amounted to nearly 4,500 GJ per day at capacity, as described by Boldt and Queneau. That translated to an approximate heat output of 2,800 GJ in hot calcine, 1,700 GJ in hot gas, and less than 1% lost through the roaster shell. Looking solely at Eqs. (1) and (2), ignoring heat from the recirculated calcine, the heat input is closer to 3,250 GJ. The hot calcine and gas was directed from the roaster to a boiler where it was then cooled. The steam generated from the boiler amounted to about 1.2 tonnes per tonne of calcine, and was sufficient for not only meeting all the plant’s steam requirements, but producing nearly enough power to meet its electrical needs, as well [4]. Fe7 S8 + 79 ̸6O2 = 1 ̸ 3Fe3 O4 + 3 Fe2 O3 + 8SO2

ð1Þ

H2 O ðliquidÞ = H2 O ðvapourÞ

ð2Þ

The proposed method for recovering iron and sulfur from pyrrhotite in this study differs from that of Inco’s Iron Ore Recovery Process, as here, its required that the nickel be removed prior to roasting. For this idealized scenario, the feed material will be a slurry of troilite and water; and the oxide product is fully hematite. Air will be supplied as the reaction gas to fluidize the bed. Equation (3) combined with Eq. (2) describe the governing reactions for this process. FeS + 7 ̸4O2 = 1 ̸ 2Fe2 O3 + SO2

ð3Þ

Following similar specification to the Inco roasting process, HSC Chemistry was employed to perform heat and mass balance for roasting troilite. Figure 1 represents this balance for 1 tonne of troilite in the feed. Per tonne of dry FeS treated, approximately 900 kg of Fe2O3 and 730 kg of SO2 are produced. The heat input is approximately 6,130 MJ per tonne pyrrhotite oxidized and the excess heat, omitting losses through the roaster shell, is about 2930 MJ (∼50% of the heat input). Scaling the input feed to match that of an Inco roaster’s daily capacity of 600 tonnes of dry troilite in a 25 weight percent water slurry, approximately 1,650 tonnes of air would need to be supplied for a dead roast to hematite, and the total heat input becomes roughly 3,700 GJ per day at capacity, with 1,750 GJ of excess energy. This energy value closely aligns to that of Inco’s roasting operation, and confirms the viability of this proposed process [5].

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Fig. 1 Theoretical heat and mass balance for roasting 1 tonne of troilite

Pyrrhotite Oxidation Kinetics Experimental Procedure Natural pyrrhotite sourced from Glencore’s Strathcona Mill was used in the following experiments. The composition of the as-received pyrrhotite powder was determined using x-ray fluorescence (XRF) analysis and is shown in Table 1. As a reference, pure pyrrhotite has a theoretical sulfur content of approximately 36–39% by weight, so it can be estimated that the purity of this natural pyrrhotite is close to 90%. In preparation, the pyrrhotite was grouped by particle size using the screening method. A Setaram 92 thermogravimetric analyzer (TGA) was used to measure the mass change of the pyrrhotite sample when exposed to oxygen at elevated temperatures as a function of time. The average sample size used was approximated 20 mg, which was measured into an alumina crucible before being placed into the furnace. Prior to heating, the TGA was vacuumed to eliminate any trapped air, and then purged with nitrogen for several minutes. Nitrogen was also used as the balance protection gas which would then flow into the reaction chamber at a rate of 10 mL/ min. The furnace was heated at a rate of 25 °C per minute in a nitrogen atmosphere. Once the target holding temperature was reached, the atmosphere was switched to an oxygen-nitrogen mix for the remaining portion of the test. Mass recording was

Na

0.42

Element

Weight percent

1.17

Mg

1.71

Al 6.35

Si 0.03

P

Table 1 Composition of natural pyrrhotite using XRF analysis 34.07

S 0.21

K 0.65

Ca 0.14

Ti 0.03

Mn

52.27

Fe

0.05

Co

1.38

Ni

0.33

Cu

0.03

Zn

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done in 4 s intervals and continued until no visible mass change was observed, implying that oxidation was complete. The atmospheric pressure inside the furnace was maintained at 101.3 kPa for the entire heating cycle. The effect of the following variables on the oxidation rate were studied in order: gas flow rate, particle size, temperature, and oxygen partial pressure. The experimental conditions are provided in Table 2.

Results and Discussion Under the experimental conditions described above, all of the 17 pyrrhotite samples exhibited mass loss as a result of oxygen being introduced to the atmosphere at these high temperatures. The results in Figs. 2, 3, 4, 5 and 6 are presented in terms of the mass loss fraction, which can be defined as (m0 − mt)/m0, where m0 is the initial sample mass at the onset of mass loss and mt is the mass of the sample at time t. 1. Effect of gas flow rate Gas flow rate was selected as the first variable to test as this variable can influence the oxidation rate by restricting mass transfer if inadequate gas is being continuously supplied to the surface (i.e. the reaction interface) [6]. The gas flow rate

Table 2 Experimental conditions of the TGA study Sample

Gas flow rate (mL/min)

Particle size (µm)

Temperature (°C)

PO2 (atm)

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16, repeat of 4 17, repeat of 4

50 100 150 200 200 200 200 200 200 200 200 200 200 200 200 200 200

38–45 38–45 38–45 38–45 45–63 63–75 75–106 106–150 38–45 38–45 38–45 38–45 38–45 38–45 38–45 38–45 38–45

850 850 850 850 850 850 850 850 700 750 800 900 850 850 850 850 850

0.21 0.21 0.21 0.21 0.21 0.21 0.21 0.21 0.21 0.21 0.21 0.21 0.10 0.15 0.97 0.21 0.21

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Fig. 2 (left to right)—Mass loss fraction of pyrrhotite versus time for different experimental variables (1) gas flow rate

Fig. 3 (left to right)—Mass loss fraction of pyrrhotite versus time for different experimental variables (2) particle size

testing began at the lowest value, which presumably would have the lowest oxidation rate, and then was increased by 50 mL/min increments until the oxidation rate no longer appeared to be impacted by the gas flow rate. These results are shown in Fig. 2. Flow rates above 200 mL/min showed a similar rate of oxidation (within the range of error) for the majority of the oxidation process. Therefore, for testing of the other three variables, a gas flow rate of 200 mL/min was used. 2. Effect of particle size To determine the effect of particle size on the reaction kinetics, five different particle size ranges were tested, as shown in Fig. 3. As expected, the highest slope, corresponding to the fasted reaction rate, was observed for the sample containing the smallest particle sizes, 38–45 µm. This is because the smaller particle sizes provide

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Fig. 4 (left to right)—Mass loss fraction of pyrrhotite versus time for different experimental variables (3) temperature

Fig. 5 (left to right)—Mass loss fraction of pyrrhotite versus time for different experimental variables (4) oxygen partial pressure

more surface area for the reaction to occur at. That being said, between 45 and 150 µm there was not a significant variance in the observed slope. Comparing this result to that of Aracena et al. [6] for oxidation of pyrite, they observed that oxidation rate decreased with increasing particle size between the range of 12.3–33 µm. It was also observed that particle size, compared to the other variables, had the most impact on the range of final mass loss fractions for the samples. There are two conceivable explanations for this. As seen in literature, the oxidation mechanism for iron sulfide particles is often described by using the Shrinking Unreacted-Core model. It describes five steps that contribute to the progression of the reaction. The second step is defined as diffusion of the gaseous reactant through the outer product layer of the particle towards the unreacted core [7, 8]. For this case, it is unlikely

Kinetics of Oxidation of Pyrrhotite

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Fig. 6 Pyrrhotite mass loss fraction as a function of time for samples 4, 16, and 17 using the following conditions: gas flow rate = 200 mL/min, particle size = 38–45 µm, temperature = 850 °C, and oxygen partial pressure = 21.3 kPa

that this is a contributing factor because conversion of FeS to Fe2O3 is accompanied with ∼20% reduction in volume, i.e. significant pore formation, which should result in improved diffusion within the oxide layer. An alternative explanation could be that there is potentially an uneven distribution of gangue material across the different particle size groups. 3. Effect of temperature Five different temperatures were used to study the impact on oxidation rate, as shown in Fig. 4. The optimal temperature appears to be 850 °C. From the temperature range of 700–850 °C, the results are in line with results for similar studies. Aracena et al. [6] also observed oxidation rate and final mass loss fraction of pyrite to have a dependence on temperature. Asaki et al. [9] studied the oxidation of dense FeS, and observed that the rate of oxidation at 850 °C was significantly faster than at 700 and 750 °C. The mass loss at 900 °C was the lowest and also had the slowest oxidation rate. As noted before, this reaction is highly exothermic, and oxidation may have caused the sample to sinter or melt, which in turn would lower the surface area and limit the progression of the reaction. This will need to be further explored by repeating testing at this temperature and analyzing the physical structure of the sample using scanning electron microscopy. 4. Effect of oxygen partial pressure The effect of oxygen partial pressure was tested by changing the O2/N2 ratio of the oxidizing gas. As expected and seen in other studies [6, 9, 10], the oxygen partial pressure has an apparent influence on the reaction rate. This variable, based on the experimental results in Fig. 5, appears to have less of an impact on the final mass loss fractions compared to the other variables tested in this study. Under the condition where the oxygen partial pressure was 97.8 kPa, the reaction rate increased drastically. Further study of the particle’s physical behavior will need to be done to

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identify if particle ignition or melting has occurred, as studies by Asaki et al. [10] and Dunn [11] for iron sulfides both identify that under vigorous oxidation conditions, such as high PO2 and temperature, such results can be observed. 5. Optimal conditions for fastest oxidation rate For the selected ranges of the four testing variables, it is determined that the fastest oxidation rate occurs when the gas flow rate is at least 200 mL/min, the particle size is between 38 and 45 µm, the temperature is at 850 °C, and the atmospheres is pure O2. For practical reasons, air is preferred so an oxygen partial pressure of 21.3 kPa was considered as optimal. In Fig. 6, the reproducibility of the results under these specific conditions are shown. The figure shows that for sample 4, 16, and 17, all under identical testing conditions, the oxidation results are highly reproducible.

Conclusions This study explored pyrometallurgical treatment of pyrrhotite for producing iron oxide and sulfur dioxide. Through materials and energy balance, it was concluded that 6,130 MJ of heat, 900 kg of hematite, and 700 kg of SO2 gas could be generated per tonne of troilite. This value, with respect to Inco’s daily roasting outputs, was in accordance with what would be expected on a commercial processing scale. In studying the optimal conditions for oxidizing natural pyrrhotite, it was determined that gas flow rate, particle size, temperature, and oxygen partial pressure all have an impact on the oxidation rate. For the variable ranges selected in this study, gas flow rate and oxygen partial pressure appeared to cause the most significant variance in oxidation rate, while particle size and temperature appeared to have the most effect on the fraction of the pyrrhotite that would convert to an oxide. It was concluded that the optimal conditions would use a gas flow rate of at least 200 mL/ min, a particle size between 38–45 µm, a temperature of 850 °C, and an oxygen partial pressure to that of air. In future studies, work will be done to confirm the exact composition of the oxide material formed and characterize the material’s microstructure as a result of roasting.

References 1. Rezaei S, Liu F, Marcuson S, Muinonen M, Lakshmanan VL, Sridhar R, Barati M (2017) Canadian pyrrhotite treatment: the history, inventory and potential for tailings processing. Can Metall Q:1–8 2. Peek E, Barnes A, Tuzun A (2011) Nickeliferous pyrrhotite-‘waste or resource?’. Miner Eng 24(7):625–637 3. Sridhar R, Dalvi A, Bakker HF, Illis A (1976) Recovery of nickel from nickeliferous pyrrhotite by a thermal upgrading process. Can Metall Q 15(3):255–262 4. Boldt JR, Queneau P (1967) The winning of nickel. Longmans Canada Ltd.

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5. Roine A (2007) HSC chemistry. Outokumpu Research Oy, Pori, Finland 6. Aracena A, Jerez O, Ortiz R, Morales J (2016) Pyrite oxidation kinetics in an oxygen-nitrogen atmosphere at temperatures from 400 to 500 °C. Can Metall Q 55(2):195–201 7. Yagi S, Kunii D (1961) Fluidized-solids reactors with continuous solids feed—II. Chem Eng Sci 16(3–4):372–379 8. Levenspiel O (1999) Chemical reaction engineering, 3rd edn. Wiley 9. Asaki Z, Matsumoto K, Tanabe T, Kondo Y (1983) Oxidation of dense iron sulfide. Metall Trans B 14(1):109–116 10. Asaki Z, Mori S, Ikeda M, Kondo Y (1985) Oxidation of pyrrhotite particles falling through a vertical tube. Metall Trans B 16(3):627–638 11. Dunn JG (1997) The oxidation of sulphide minerals. Thermochim Acta 300:127–139

Formation Mechanism of Ferronickel Alloy Due to the Reaction Between Iron and Nickeliferous Pyrrhotite at 850–900 °C Feng Liu, Mansoor Barati and Sam Marcuson

Abstract A thermal upgrading process by which nickel value can be concentrated in a ferronickel alloy is a possible alternative to treat Sudbury pyrrhotite (Po) tailings with nickel content of 0.5–1.5 wt%. The basis of this process is precipitation of Ni from Po at high temperature once Fe/S ratio in the iron-deficient Po is shifted towards stoichiometric or near stoichiometric FeS (troilite) either by the addition of iron and/or the removal of sulfur. For the iron addition route, the reaction between elemental iron and nickeliferous pyrrhotite to produce ferronickel alloy and Ni-depleted iron sulfide phase plays a critical rule. In this paper, the formation mechanism of ferronickel alloy was investigated using the diffusion couple technique to better understand the nickel diffusion behavior in the iron and sulfide phases. Keywords Nickeliferous pyrrhotite Sulfur vapor Nickel diffusion





Thermal upgrading



Ferronickel alloy

Introduction Over the past 50 years, the nickel mining operations in Sudbury, Canada have generated nickeliferous pyrrhotite tailings amounting to 75–100 Mt on a dry basis [1]. The nickeliferous pyrrhotite found in Sudbury is often reported to contain 0.6– 0.95%Ni in solid solution [2–7] with an average content of 0.7% [8]. Also, due to the presence of micron-sized pentlandite intergrowths (1–5 μm) embedded in the pyrrhotite matrix, as well as free pentlandite particles which inadvertently report to the pyrrhotite tailings during the beneficiation of nickel ores, Sudbury pyrrhotite tailings may contain up to 1.5%Ni, depending on the orebody and mill operation F. Liu (✉) ⋅ M. Barati University of Toronto, 184 College Street, Toronto, ON M5S 3E4, Canada e-mail: [email protected] S. Marcuson Independent Consultant, Mississauga, ON, Canada © The Minerals, Metals & Materials Society 2018 B. Davis et al. (eds.), Extraction 2018, The Minerals, Metals & Materials Series, https://doi.org/10.1007/978-3-319-95022-8_33

415

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conditions [2]. In general, these tailings are characterized by a nickel deportment at least 50% in pyrrhotite (in solid solution) and the other half in pentlandite [9–11]. At an average 0.8%Ni content including pentlandite, the pyrrhotite tailings accumulated to date in the Sudbury region are estimated to contain 600–800 kt Ni. In addition, at present, the two mills in Sudbury, Strathcona mill and Clarabelle mill, have the capacity to produce fresh pyrrhotite tailings at a rate of approximately 1500 tpd [12] and 4300 tpd [8, 13], respectively. If a processing plant is built to treat these tailings, the potential revenue from this Ni resource would be over 4 billion USD, assuming 70% plant recovery and a Ni price of 9500 USD/t [1]. Although commercial operations of the past recovered the nickel value from this pyrrhotite [14–16], currently it is not processed for value recovery. Instead, the pyrrhotite tailings rejected from the mills plants are dumped into tailings ponds, and then covered with layers of low-sulfur rock tails and water. Such deposition of pyrrhotite tailings is intended to reduce the environmental risk associated with acid mine drainage (AMD) caused by oxidation when pyrrhotite is exposed to air and moisture [12, 17, 18]. However, storage of these tailings in ponds may not be viewed as a long-term solution due to the maintenance costs and ever-present risk of AMD. In consideration of the significant amounts of historical pyrrhotite tailings that could be readily reclaimed via surface mining, and the fresh finely-ground tailings still being generated, it is desirable that an integrated, energy-efficient, cost-effective and environmentally-sound process should be developed to treat the pyrrhotite tailings as a resource for extraction of nickel rather than a waste. A thermal upgrading process to concentrate the nickel value in a metallic alloy phase followed by magnetic separation was first studied by INCO in 1970s [19, 20]. The basic principle of this process can be illustrated using the Fe–Ni–S phase diagram. As shown in Fig. 1, at 900 °C a mono-sulfide solid solution (mss) exists between Fe1−xS and Ni1−xS. Also, it is noted that the Fe-rich portion of this ternary system at 900 °C is dominated by a two-phase assemblage ‘γ + mss’ where γ denotes the ferronickel alloy with a FCC crystal structure. At 900 °C, the typical composition of Sudbury nickeliferous pyrrhotite is located within the area of mss, close to the Fe–S side. The thermal upgrading process can be viewed as the change in phase associations from one-phase region ‘mss’ to two-phase region or even three-phase region ‘alloy + mss’ as a result of a change in the Fe/S ratio. This can be achieved by adding iron into and/or removing sulfur from the pyrrhotite matrix, thereby converting nickeliferous pyrrhotite to a mixture of strongly magnetic alloy phase and weakly magnetic sulfide phase. Most of the nickel will be concentrated in the alloy phase, whereas the weakly magnetic fraction, consisting mainly of stoichiometric or nearly stoichiometric FeS is found to contain less than 0.5%Ni [21]. In other words, the possibility of extracting nickel by the thermal upgrading process relies upon the difference of its equilibrium concentration in mss versus γ, which can be determined from the position of tie lines in the two-phase region ‘γ + mss’. Therefore, the main task of the thermal upgrading process of nickeliferous pyrrhotite for extracting nickel is to increase the Fe/S molar ratio, approaching unity. One way to increase the Fe/S ratio is to add elemental iron into this ternary system. Under an inert atmosphere at high temperatures, the diffusion of Fe, Ni, and S will

Formation Mechanism of Ferronickel Alloy due to the Reaction …

417

Fig. 1 Phase association of Fe–Ni–S system at 900 °C and illustration of the basic principle of thermal upgrading process (Calculated with FactSage 7.0 [22])

shift the composition to a new equilibrium state corresponding to the modified composition. The present study employs a diffusion couple technique to investigate the interactions between Fe and Po. The knowledge of the reaction characteristics will be used to build the model depicting the formation of ferronickel alloy.

Experimental Materials and Diffusion Couple Preparation Samples of pyrrhotite tailings were received from the Strathcona mill (Glencore). To improve the grade in the pyrrhotite phase, magnetic separation using a Davis tube was performed to collect monoclinic pyrrhotite particles. Table 1 lists the chemical composition of as-received pyrrhotite tailings as well as their magnetic concentrate. Compared with the theoretical composition of Fe1−xS (x = 0–0.2) [23] the major component of the magnetic concentrate is confirmed to be the pyrrhotite phase (>98% in weight) with nickel in solid solution. Particle size distribution of the as-received pyrrhotite tailings and their magnetic concentrate was determined using a laser scattering particle size distribution analyzer. As seen in Fig. 2, the particle size range of the magnetic concentrate is narrower compared to that of the as-received pyrrhotite tailings. Both large and fine Table 1 Chemical composition of as-received pyrrhotite tailings and their magnetic concentrate, wt% Elements

Fe

Ni

S

Cu

Others

As-received Po tailings Magnetic Po concentrate

47.0 61.0

1.02 0.83

29.5 36.2

0.27 0.05

22.2 1.92

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Fig. 2 Particle size distribution: a as-received pyrrhotite tailings; b magnetic concentrate

particles in the as-received pyrrhotite tailings which are either silicate minerals or pentlandite were rejected during the magnetic separation. Moreover, for the as-received pyrrhotite tailings a separate peak centered at around 45 μm can be identified (bimodal distribution). Based on the particle size distributions of these two materials, the magnetic concentrate was screened to produce particles measuring 38–45 μm in diameter. Electron probe micro analysis (EPMA) was then performed to determine the chemical composition of these particles. The average composition of 10 points is shown in Table 2. For comparison, the result of probe analysis of monoclinic pyrrhotite sampled from the Sudbury orebody is also shown in Table 2 [4]. Two types of diffusion couples were employed in this study, as shown in Fig. 3. For both types, the screened magnetic particles were used on the pyrrhotite side. On the iron side, either a bulk, dense sheet of iron or a compacted layer of iron powder were used. The diffusion couple was placed in an alumina crucible, and heat treated.

Experimental Setup and Procedure A Setaram TG92 was used to heat treat all of the samples; a schematic diagram is shown in Fig. 4. Before each test, the whole system was evacuated twice to remove any residual oxygen. Then the sample was heat treated to the target temperature (850–900 °C) at 25 °C/min under argon atmosphere, soaked for a predetermined period of time, and left inside the furnace to cool down at 25 °C/min. After the thermal treatment step, the sample together with the crucible was mounted in epoxy, ground, polished, and analyzed by EPMA. Table 2 Average chemical composition of the pyrrhotite phase determined by EPMA, wt%

Elements

Fe

Ni

S

Co

Magnetic Po concentrate Monoclinic Po [4]

59.4 59.5

0.79 0.82

39.7 39.7

0.07 –

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419

Fig. 3 Schematic illustration of diffusion couples with different iron materials: a dense, bulk iron; b particulate iron

Fig. 4 Schematic diagram of Setaram TG92 and thermal treatment schedule [24]

Results and Discussion Figure 5 shows a micrograph of a diffusion couple specimen heat treated at 850 °C for 1.5 h. This diffusion couple consisted of a piece of dense iron foil embedded in the magnetic pyrrhotite concentrate. After thermal treatment, the most obvious change was the drastic reduction in the thickness of the iron foil from an initial thickness of 100 μm to about 80 μm. Adjacent to the remaining iron, a reaction layer composed mainly of near stoichiometric FeS was found. This newly-formed iron sulfide layer features a dense morphology which is more similar to the iron

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Alloy Composition, wt% Fe

Ni

S

1

84.4

14.5

0.86

2

82.2

17.3

0.06

Fig. 5 Micrograph of a cross-section of the heat-treated diffusion couple consisting of 100 μm thick iron foil embedded in the magnetic pyrrhotite powder compact (850 °C, 1.5 h)

layer, and very different from that of the original pyrrhotite layer. The formation of this reaction layer must be caused by sulfur transport from the sulfide phase to the surface of the iron foil, and subsequent reaction with the elemental iron. Condit et al. [25] discussed the growth mechanism of ferrous sulfide as a result of Fe–S reaction, and verified the importance of iron diffusion and sulfur vapor transport. Since the mobility of sulfide ions in the pyrrhotite phase is far lower than that of iron at temperatures higher than 347 °C [25], sulfur vapor transport toward the iron phase is thus considered as the major mechanism whereby the dense iron sulfide forms. Considering that the temperature used in the current study was not high enough to cause either the formation of a liquid phase or extensive sintering, the long-distance iron diffusion within the sulfide phase should be limited. Due to the non-stoichiometry of pyrrhotite, the fugacity of S2 versus temperature when the pyrrhotite composition is varied can be described by the following equation derived by Toulmin et al. [26] log10 fS2 = ð70.03 − 85.83NFeS Þð1000 K ̸ T − 1Þ + 39.30

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 − 0.9981NFeS − 11.91

where fS2 is the fugacity of S2 relative to the ideal diatomic gas at 1 atm, NFeS which is the mole fraction of FeS in the system FeS–S2 (Fe1−xS is considered to consist of FeS and S2) can be calculated as twice the atomic fraction of iron for a given Fe1 −xS, and T is the absolute temperature with unit Kevin. Figure 6 shows the plots of fS2 as a function of temperature in the Fe–S system. The above equation was used to derive the pyrrhotite field where five different pyrrhotite polymorphs with different NFeS were considered. The pyrrhotite-pyrite curve was determined by employing the data measured by Toulmin et al. [26], and the boundary between liquid sulfur and S2 vapor was calculated using FactSage 7.0 [22]. It can be seen that at 850 °C, the fugacity of S2 in equilibrium with monoclinic Fe7S8 (NFeS = 0.933) is about 0.3 atm. This significant sulfur vapor pressure would contribute to the transport of sulfur to the iron foil.

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421

Fig. 6 Temperature dependence of the fugacity of S2 in the Fe–S system

Further inspection of the newly-formed iron sulfide shows the presence of tiny metallic particles measuring a few micrometers (

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