VOLUME 114
NO. 4
APRIL 2014
CENTRIFUGAL & MECHANICAL INDUSTRIES
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THE AFRICA PRIZE FOR ENGINEERING INNOVATION INNOVA ATION T FOR
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The Royal Academy of Engineering announces the launch of the Africa Prize for Engineering Innovation. The prize aims to stimulate, celebrate and reward innovation and entrepreneurship in sub-Saharan Africa. Applications are invited from engineers affiliated with a university or research institution in sub-Saharan Africa who have developed innovations that can provide scalable solutions to local challenges. Engineers from sub-Saharan Africa and from all engineering disciplines are eligible to enter.
Shortlisted applicants will receive a six-month package of tailored support to help them achieve commercial success from their innovation. The overall winner of the GBP25,000 Prize will be announced at a high-profile event in May 2015. The Africa Prize is generously supported by
Applications close at midnight on 30 May 2014.
The Shell Centenary Scholarship Fund Consolidated Contractors Company ConocoPhillips Mo Ibrahim Foundation
Find out more at: www.raeng.org.uk/AfricaPrize or africaprize@raeng.org.uk
The Southern African Institute of Mining and Metallurgy OFFICE BEARERS AND COUNCIL FOR THE 2013/2014 SESSION Honorary President Mark Cutifani President, Chamber of Mines of South Africa Honorary Vice-Presidents Susan Shabangu Minister of Mineral Resources, South Africa Rob Davies Minister of Trade and Industry, South Africa Derek Hanekom Minister of Science and Technology, South Africa President M. Dworzanowski President Elect J.L. Porter Vice-Presidents R.T. Jones C. Musingwini Immediate Past President G.L. Smith Honorary Treasurer J.L. Porter Ordinary Members on Council H. Bartlett N.G.C. Blackham V.G. Duke M.F. Handley W. Joughin A.S. Macfarlane D.D. Munro
S. Ndlovu G. Njowa S. Rupprecht A.G. Smith M.H. Solomon D. Tudor D.J. van Niekerk
Past Presidents Serving on Council N.A. Barcza R.D. Beck J.A. Cruise J.R. Dixon F.M.G. Egerton A.M. Garbers-Craig G.V.R. Landman
R.P. Mohring J.C. Ngoma R.G.B. Pickering S.J. Ramokgopa M.H. Rogers J.N. van der Merwe W.H. van Niekerk
Branch Chairmen DRC
S. Maleba
Johannesburg
I. Ashmole
Namibia
G. Ockhuizen
Pretoria
N. Naude
Western Cape
T. Ojumu
Zambia
H. Zimba
Zimbabwe
S.A. Gaihai
Zululand
C. Mienie
PAST PRESIDENTS *Deceased * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * *
W. Bettel (1894–1895) A.F. Crosse (1895–1896) W.R. Feldtmann (1896–1897) C. Butters (1897–1898) J. Loevy (1898–1899) J.R. Williams (1899–1903) S.H. Pearce (1903–1904) W.A. Caldecott (1904–1905) W. Cullen (1905–1906) E.H. Johnson (1906–1907) J. Yates (1907–1908) R.G. Bevington (1908–1909) A. McA. Johnston (1909–1910) J. Moir (1910–1911) C.B. Saner (1911–1912) W.R. Dowling (1912–1913) A. Richardson (1913–1914) G.H. Stanley (1914–1915) J.E. Thomas (1915–1916) J.A. Wilkinson (1916–1917) G. Hildick-Smith (1917–1918) H.S. Meyer (1918–1919) J. Gray (1919–1920) J. Chilton (1920–1921) F. Wartenweiler (1921–1922) G.A. Watermeyer (1922–1923) F.W. Watson (1923–1924) C.J. Gray (1924–1925) H.A. White (1925–1926) H.R. Adam (1926–1927) Sir Robert Kotze (1927–1928) J.A. Woodburn (1928–1929) H. Pirow (1929–1930) J. Henderson (1930–1931) A. King (1931–1932) V. Nimmo-Dewar (1932–1933) P.N. Lategan (1933–1934) E.C. Ranson (1934–1935) R.A. Flugge-De-Smidt (1935–1936) T.K. Prentice (1936–1937) R.S.G. Stokes (1937–1938) P.E. Hall (1938–1939) E.H.A. Joseph (1939–1940) J.H. Dobson (1940–1941) Theo Meyer (1941–1942) John V. Muller (1942–1943) C. Biccard Jeppe (1943–1944) P.J. Louis Bok (1944–1945) J.T. McIntyre (1945–1946) M. Falcon (1946–1947) A. Clemens (1947–1948) F.G. Hill (1948–1949) O.A.E. Jackson (1949–1950) W.E. Gooday (1950–1951) C.J. Irving (1951–1952) D.D. Stitt (1952–1953) M.C.G. Meyer (1953–1954)
* * * * * * * * * * * * * * * * * * * * * * * *
*
*
*
*
*
L.A. Bushell (1954–1955) H. Britten (1955–1956) Wm. Bleloch (1956–1957) H. Simon (1957–1958) M. Barcza (1958–1959) R.J. Adamson (1959–1960) W.S. Findlay (1960–1961) D.G. Maxwell (1961–1962) J. de V. Lambrechts (1962–1963) J.F. Reid (1963–1964) D.M. Jamieson (1964–1965) H.E. Cross (1965–1966) D. Gordon Jones (1966–1967) P. Lambooy (1967–1968) R.C.J. Goode (1968–1969) J.K.E. Douglas (1969–1970) V.C. Robinson (1970–1971) D.D. Howat (1971–1972) J.P. Hugo (1972–1973) P.W.J. van Rensburg (1973–1974) R.P. Plewman (1974–1975) R.E. Robinson (1975–1976) M.D.G. Salamon (1976–1977) P.A. Von Wielligh (1977–1978) M.G. Atmore (1978–1979) D.A. Viljoen (1979–1980) P.R. Jochens (1980–1981) G.Y. Nisbet (1981–1982) A.N. Brown (1982–1983) R.P. King (1983–1984) J.D. Austin (1984–1985) H.E. James (1985–1986) H. Wagner (1986–1987) B.C. Alberts (1987–1988) C.E. Fivaz (1988–1989) O.K.H. Steffen (1989–1990) H.G. Mosenthal (1990–1991) R.D. Beck (1991–1992) J.P. Hoffman (1992–1993) H. Scott-Russell (1993–1994) J.A. Cruise (1994–1995) D.A.J. Ross-Watt (1995–1996) N.A. Barcza (1996–1997) R.P. Mohring (1997–1998) J.R. Dixon (1998–1999) M.H. Rogers (1999–2000) L.A. Cramer (2000–2001) A.A.B. Douglas (2001–2002) S.J. Ramokgopa (2002-2003) T.R. Stacey (2003–2004) F.M.G. Egerton (2004–2005) W.H. van Niekerk (2005–2006) R.P.H. Willis (2006–2007) R.G.B. Pickering (2007–2008) A.M. Garbers-Craig (2008–2009) J.C. Ngoma (2009–2010) G.V.R. Landman (2010–2011) J.N. van der Merwe (2011–2012)
Honorary Legal Advisers Van Hulsteyns Attorneys
Corresponding Members of Council Australia:
I.J. Corrans, R.J. Dippenaar, A. Croll, C. Workman-Davies
Auditors Messrs R.H. Kitching
Austria:
H. Wagner
Secretaries
Botswana:
S.D. Williams
Brazil:
F.M.C. da Cruz Vieira
China:
R. Oppermann
The Southern African Institute of Mining and Metallurgy Fifth Floor, Chamber of Mines Building 5 Hollard Street, Johannesburg 2001 P.O. Box 61127, Marshalltown 2107 Telephone (011) 834-1273/7 Fax (011) 838-5923 or (011) 833-8156 E-mail: journal@saimm.co.za
United Kingdom: J.J.L. Cilliers, N.A. Barcza, H. Potgieter USA:
J-M.M. Rendu, P.C. Pistorius
Zambia:
J.A. van Huyssteen
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APRIL 2014
The Journal of The Southern African Institute of Mining and Metallurgy
Editorial Board R.D. Beck J. Beukes P. den Hoed M. Dworzanowski M.F. Handley R.T. Jones W.C. Joughin J.A. Luckmann C. Musingwini R.E. Robinson T.R. Stacey
VOLUME 114
NO. 4
APRIL 2014
STUDENT EDITION
Editorial Consultant D. Tudor
Typeset and Published by
Printed by Camera Press, Johannesburg
Advertising Representative Barbara Spence Avenue Advertising Telephone (011) 463-7940 E-mail: barbara@avenue.co.za The Secretariat The Southern African Institute of Mining and Metallurgy ISSN 2225-6253
THE INSTITUTE, AS A BODY, IS NOT RESPONSIBLE FOR THE STATEMENTS AND OPINIONS A DVA NCED IN A NY OF ITS PUBLICATIONS. Copyright© 1978 by The Southern African Institute of Mining and Metallurgy. All rights reserved. Multiple copying of the contents of this publication or parts thereof without permission is in breach of copyright, but permission is hereby given for the copying of titles and abstracts of papers and names of authors. Permission to copy illustrations and short extracts from the text of individual contributions is usually given upon written application to the Institute, provided that the source (and where appropriate, the copyright) is acknowledged. Apart from any fair dealing for the purposes of review or criticism under The Copyright Act no. 98, 1978, Section 12, of the Republic of South Africa, a single copy of an article may be supplied by a library for the purposes of research or private study. No part of this publication may be reproduced, stored in a retrieval system, or transmitted in any form or by any means without the prior permission of the publishers. Multiple copying of the contents of the publication without permission is always illegal. U.S. Copyright Law applicable to users In the U.S.A. The appearance of the statement of copyright at the bottom of the first page of an article appearing in this journal indicates that the copyright holder consents to the making of copies of the article for personal or internal use. This consent is given on condition that the copier pays the stated fee for each copy of a paper beyond that permitted by Section 107 or 108 of the U.S. Copyright Law. The fee is to be paid through the Copyright Clearance Center, Inc., Operations Center, P.O. Box 765, Schenectady, New York 12301, U.S.A. This consent does not extend to other kinds of copying, such as copying for general distribution, for advertising or promotional purposes, for creating new collective works, or for resale.
Contents Journal Comment by H. Phillips . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . President’s Corner by M. Dworzanowski . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
v vii
Special Articles Book review—Digging Deep by N. Mayer . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . MQA gives Wits University over R20 million by K. Foss . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Manager: Regional Development . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . The SAIMM Library . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
viii ix x 346
PAPERS A critical evaluation of haul truck tyre performance and management system at Rössing Uranium Mine by T.S. Kagogo . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . A comparative study between shuttle cars and battery haulers by W.H. Holtzhausen . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Investigation of cavity formation in lump coal in the context of underground coal gasification by C. Hsu, P.T. Davies, N.J. Wagner, and S. Kauchali . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Arnot’s readiness to prevent a Pike River disaster by R. Weber. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Laser cladding AA2014 with a Al-Cu-Si compound for increased wear resistance by K.J. Kruger and M. du Toit . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Development of a method for evaluating raw materials for use in iron ore sinter in terms of lime assimilation by W. Ferreira, R. Cromarty, and J. de Villiers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . The recovery of manganese products from ferromanganese slag using a hydrometallurgical route by S.J. Baumgartner and D.R. Groot . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Wear of magnesia-chrome refractory bricks as a function of matte temperature by M. Lange, A.M. Garbers-Craig, and R. Cromarty. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Comparing the extent of the dissolution of copper-cobalt ores from the DRC Region using sulphuric acid in the presence of hydrogen peroxide and in tartaric acid by S. Stuurman, S. Ndlovu, and V. Sibanda . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
293 299 305 311 317
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VOLUME 114
NO. 4
APRIL 2014
International Advisory Board R. Dimitrakopoulos, McGill University, Canada D. Dreisinger, University of British Columbia, Canada E. Esterhuizen, NIOSH Research Organization, USA H. Mitri, McGill University, Canada M.J. Nicol, Murdoch University, Australia H. Potgieter, Manchester Metropolitan University, United Kingdom E. Topal, Curtin University, Australia
The Journal of The Southern African Institute of Mining and Metallurgy
APRIL 2014
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The Southern African Institute of Mining and Metallurgy P.O. Box 61127 Marshalltown 2107 Telephone (011) 834-1273/7 Fax (011) 838-5923 E-mail: journal@saimm.co.za
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Journal Comment
The Journal of The Southern African Institute of Mining and Metallurgy
eventually lost its status as a division of CSIR, and South Africa lost its status as the leading country for mining research. Only the coal mining sector has sustained coherent research activities, through its collaborative research programme, Coaltech. It must be acknowledged that individual mining companies do undertake and sponsor research but this confidential research is often piecemeal, with contracts awarded worldwide to institutions and individuals who are experts in particular research areas. The result of this has been a dearth of opportunities for young people in South Africa to develop their skills in mining research under the mentorship of experienced researchers. Over the past 15 years or so the focus of most mining companies has been on meeting the targets for transformation and sustainable development. Health and safety of the workforce, care for the environment, and engagement with communities are undoubtedly a vital part of mining. However, the need to develop new mining methods and technology to mine safely and economically requires hard-core mining engineering research. Since the completion of major programmes such as Deep-mine and Future-mine little has been done to consolidate the findings of these programmes and put them into practice. The future of many mines is to mine deeper and this, coupled with a significant change in the cost of employment of the underground workforce, cries out for efficient and effective mechanization, as the forerunner of automation of many mining operations. At present there is no organization or institution commanding sufficient respect from the mining industry to be the leader or custodian of the necessary research. Should it be a government department that initiates a revival of mining engineering research? Should it be the Chamber of Mines, or the CSIR, or indeed a consortium of universities? Time will tell, but time is also running out and the store of knowledge from previous research is dissipating fast. What lends optimism to the current situation is the quality of the research undertaken by the students whose work appears in this edition of the journal. The two mining papers, three mineral processing papers, and three material science papers are of a high standard and give a clear indication of the talent available to our industry. Let’s hope these young people find careers that utilize their talents in meeting the challenges faced in the mining and beneficiation of our mineral resources.
H. Phillips
APRIL 2014
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The papers in this edition of the Journal are authored or co-authored by recent graduates in mining and metallurgy. They are based on final year undergraduate projects and were presented at the annual Southern African Institute of Mining and Metallurgy’s Student Colloquium in November 2013. This was held at the University of Johannesburg, and for the first time a student from Namibia presented a paper. While the tone of the presentations ranged from outrageously flamboyant to virtual stage fright, it was abundantly clear that the research being undertaken by these students was of a remarkably high standard. The diversity of topics, together with the differences in approach by the different disciplines, was catered for by parallel sessions but delegates were unanimous in their praise of the quality of the presentations. The choice as to which ten papers should be published in the Journal must have been extremely difficult. Rather than attempting to summarize the papers or comment on them individually, I would prefer to turn to the vexing question of mining research in South Africa. I exclude metallurgy, metallurgical, and chemical engineering not due to any prejudice, but due to my ignorance of those subjects and the feeling I have that research in these areas has fared better in recent years than has mining research. This feeling is reinforced by the very recent news that the CSIR has decided to disaggregate (their word, not mine) their Centre for Mining Innovation and to reassign its remaining researchers to areas where similar competencies exist but which service many different industrial sectors and clients. This terminates 50 years of mining research on the Auckland Park site, since it was in 1964 that the Transvaal and Orange Free State Chamber of Mines formed the Chamber of Mines Research Organisation (COMRO) and established it in Carlow Road. The need for a mining research organization was recognized following the inquiry into the Coalbrook disaster, which found there was no scientific basis for the design of coal pillars and highlighted the need for systematic research. Three existing laboratories (Dust and Ventilation, Applied Physiology, and Biological and Chemical Research) were incorporated into the newly established COMRO and both a Mining Research Laboratory and a Physical Sciences Laboratory were created. An Environmental Services Division was added and in 1966 a Colliery Research Laboratory completed the Organization. During the 1970s and 1980s South Africa was undoubtedly the world leader in most aspects of mining research and the research output was prodigious. When I had the privilege of spending six months of sabbatical leave working at COMRO in 1981 there were nearly 700 employees, with three quarters of them being active researchers. The early 1990s saw structural changes and a reduction in size and in 1993 COMRO was taken over by the CSIR and became Miningtek. The staffing compliment was significantly reduced over the years and Miningtek
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ne of the key drivers for a successful South African mining industry is a pipeline of mining and metallurgy graduates. Without these individuals the sustainability of the industry will be in jeopardy. Therefore, a key aspect of SAIMM activities is the support of mining and metallurgy students and the tertiary institutions that provide their education. While there is some financial support available via the SAIMM scholarship trust fund, the type of support that the SAIMM provides is to motivate students with their studies and improve the probability of them being employed within the industry. The annual SAIMM Student Colloquium has proved to be a great success in bringing students together from different tertiary institutions. The presentations given on the mining and metallurgy projects being undertaken clearly show the quality of students being graduated. The 2013 subject matter looked at different commodities, with some of the work conducted directly with a mining operation. The topics, which included different aspects of ore transport for underground and open pit operations, material wear properties, sintering for iron production, hydrometallurgical treatment of ferromanganese slag, wear rate of refractory bricks in furnaces, and hydrometallurgical treatment of copper/cobalt ore, highlight the variety of work being conducted and its clear applicability to the mining industry. The SAIMM Student Colloquium has started to broaden the involvement of students from our southern African branches, which is an important initiative going forward. Setting a date for the colloquium that can encompass as many of the tertiary institutions as possible has always been a challenge and will continue to be, but our organizing committee always manages to get the best possible results each year and all credit to them. There is a tendency to forget that the SAIMM Western Cape branch encompasses the tertiary institutions in the Western Cape which generate many high-quality metallurgy graduates. They also host an annual event, a two day Mineral Processing conference, which highlights the significant amount of work which the undergraduates and postgraduates there produce. Last year I was asked by the Branch to give a keynote address at this annual conference, which allowed me to witness the significant contribution these tertiary institutions make towards metallurgy research and the output of metallurgy graduates. The SAIMM has also started organizing career days for mining and metallurgy students where the students have the opportunity to ask a panel of mining engineers and metallurgists questions about different career paths within the mining and metallurgy disciplines. The first event, held at Sci Bono last year, proved to be a huge success. I was on the panel, so I can testify to this. After students complete their studies and start a career in the mining industry, they become Young Professionals. This year the SAIMM hosted the first Young Professionals conference, where experiences were shared across different commodities and disciplines by quite a diverse group of young mining engineers and metallurgists. This is another example of support to students as they leave their studies and embark upon a career in the mining industry. Guidance given early in their careers will increase the probability of them staying with a career in the mining industry. The above points illustrate the extent to which the SAIMM supports mining and metallurgy students and their progression to Young Professionals. I cannot overemphasise how important this support is, and how essential it is for the SAIMM to examine ways to expand this support.
tʼs iden s e r P er Corn
The Journal of The Southern African Institute of Mining and Metallurgy
APRIL 2014
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M. Dworzanowski President, SAIMM
Book Review—Digging Deep
J
ade Davenport, the author of Digging Deep, has written a very readable book on the historic contribution that mining has made in developing South Africa into a modern industrial state. It is not necessarily a must-read for people in the mining industry, but the book is well written and should be seen as a South African mining biography. Digging Deep can be regarded as an anthology, each component of which can be read in isolation, but with a thread that links the parts into the complete story of the vital role that the mining industry played in this country’s economic and political history. I would, however, recommend that the chapters be read in sequence. The book will be of interest to a broad spectrum of South African readers, as it is not a book on the technical aspects of the industry. There is, however, sufficient detail that will allow mining professionals to find it a satisfying read. The history is clearly well researched and is detailed enough to help the reader get a good perspective of the social, political, and economic impact on the development of South Africa. In essence, it should have general appeal. Although replete with the names of key players in the history of mining and details of production and costs, the book remains an easy read. However, the terminology may be unfamiliar to the general reader, and footnotes on the relevant pages would have been helpful. The author has tried to use the relevant terms in the historical context of currency of the day and imperial/metric systems, but has deviated occasionally in parts of the book. In an early chapter, an attempt is made to equate the value of money then in use to the current rand. The author is to be commended for writing the book in such a way that each chapter’s history is complete, yet linked along the historic path to the other chapters. Digging Deep, by Jade Davenport, is published by Jonathan Ball.
The reviewer Nap Mayer was involved in the mining industry from 1960 to 2000 with experience in copper, diamonds, coal, and gold. As Managing Director of Anglo American’s Gold and Uranium Division, he served on the Anglo American Board as an Alternate Director. He also served on the Atomic Energy Board as Vice- Chairman for a three-year period from 1996. Nap is a retired, corporate member of the SAIMM.
N. Mayer
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The Journal of The Southern African Institute of Mining and Metallurgy
MQA gives Wits University over R20 million
The Journal of The Southern African Institute of Mining and Metallurgy
Prof Frederick Cawood, head of the Wits School of Mining Engineering, Mr Yunus Omar, CFO of the MQA and Vice-chancellor and principle of Wits University, Prof Adam Habib
R100 000 will be used to support students whose studies are being affected because they can’t afford necessities such as spectacles. The MQA also supports the kitchen project, which feeds students who can’t afford lunch. Cawood said the School had seen an almost 99% success rate in students who had been assisted in this manner. The CFO of the MQA, Yunus Omar, told student and staff representatives from Wits that the MQA was comprised of people who had been in their shoes. ‘They know what the students and lecturers are going through’, he said.
K. Foss
APRIL 2014
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17 March 2014 - Johannesburg: The Mining Qualifications Authority (MQA) handed over a cheque for more than R20 million to Wits University on Friday, 14 March 2014. The money will go towards support for seven lecturers in mining engineering and bursaries for 236 students in the following disciplines: analytical, chemical, electrical, industrial, mechanical, metallurgical, and mining engineering, and geology. The Head of the School of Mining Engineering, Professor Fred Cawood, said the long-standing partnership between Wits and the MQA dated back to 2005 and had strengthened to the point where it was valued at such a significant sum of money. ‘This commitment speaks volumes about the MQA and sets an example for other SETAs’, he said. Vice-Chancellor and Principal of Wits, Professor Adam Habib, echoed Cawood’s opinion that the MQA had set an example for other industries, and reminded those present that the historical disenfranchisement of some South Africans had created enormous levels of inequality that could only be addressed through collective action. ‘The VC can no longer say that his responsibilities end at the gates of the university. The CEO can no longer say that his responsibilities end with the company’s shareholders. How we begin to bridge institutional boundaries has become important. This partnership with the MQA is testimony to what can be done’, said Habib. The total amount of the partnership is R23 592 113.03. The total amount of support for lecturers is R4 624 113.03, and the total amount to be given in bursaries is R18 868 000.00. Habib said the bursaries, to be given to disadvantaged students, would send a powerful message of hope to the poor that talented people have access to one of the best universities in the country, and that the support that would be given to lecturers was an investment in the creation of a new black professoriate.
The Southern African Institute of Mining and Metallurgy (SAIMM) MANAGER: REGIONAL DEVELOPMENT Applications are invited for the newly created position of Manager: Regional Development with the Southern African Institute of Mining and Metallurgy. The primary purpose of this job is to sustainably expand non-South African membership of the SAIMM by establishing and supporting National Branches in neighbouring countries. Secondary activities are related to representation of the SAIMM on appropriate industry-level forums that support membership growth and Institute sustainability.
Requirements of Position The incumbent will be responsible for, amongst other duties, the development and execution of a regional marketing and growth strategy that is aligned with SAIMM strategy and objectives; the development and support of new and existing National Branches across Southern Africa; provision of direct strategic and tactical support to the National Branches; co-ordination of SAIMM support across National Branches to achieve an optimal Institute outcome; as well as the development and implementation of National Branch events calendars that build stakeholder engagement and ensure Branch sustainability. Other duties will include the establishment of annual operating plans and budgets within a three-year strategic plan for each National Branch. The person appointed must be able to engage at a senior level in the minerals industry, and to communicate on technical matters in forums representing the Professional membership. Generic marketing skills around monitoring and analysing market trends, as well as studying competitors' products and services, are some of the marketing skills required. Other qualities include the ability to respond well to pressure; creative thinking; good presentation skills; and the ability to motivate and lead teams. Significant travel is required to build and maintain the National Branches.
Experience The successful applicant should have a minerals industry background with 10+ years minerals industry experience; broad knowledge of the global mining industry from resource to customer; and be able to demonstrate people leadership experience i.e. has held a substantive position and/or project management experience. It is recommended, but not a requirement, that the incumbent is a current Fellow/Member of the SAIMM or other professional entity. Tertiary qualifications related to marketing and the minerals industry would be beneficial.
Enquiries: Sam Moolla; Tel: +27 11 834-1273/7; Email: sam@saimm.co.za. To apply: Please submit a letter of motivation and a detailed CV with the names, contact numbers, and e-mail addresses of three references. Closing date: 30 June 2014.
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The Journal of The Southern African Institute of Mining and Metallurgy
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Student Papers A critical evaluation of haul truck tyre performance and management system at Rössing Uranium Mine by T.S. Kagogo . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 293 The factors affecting haul truck tyre performance and the effectiveness of the management system in the load-and-haul operation are investigated with the aim of increasing tyre performance in terms of running time to failure A comparative study between shuttle cars and battery haulers by W.H. Holtzhausen. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 299 Two different underground batch coal haulers are compared to determine the more viable machine to implement. Running costs, capital costs, and maintenance costs are compared, and production rates and availability are used to determine which type of machine would be more reliable. Investigation of cavity formation in lump coal in the context of underground coal gasification by C. Hsu, P.T. Davies, N.J. Wagner, and S. Kauchali . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 305 This project investigated, on a laboratory scale, cavity formation within a coal block due to the combustion reactions that take place during the coal gasification process in an underground environment. The results are compared with earlier work and a similar trend was observed, despite a slightly different methodology being employed. Arnot’s readiness to prevent a Pike River disaster by R. Weber . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 311 The effectiveness of Arnot Colliery’s measures to prevent conditions favourable for methane explosions is evaluated, with particular reference to the causes of the Pike River disaster in New Zealand. Laser cladding AA2014 with a Al-Cu-Si compound for increased wear resistance by K.J. Kruger and M. du Toit . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 317 Aluminium alloy 2014 (AA2014) was coated with a 1.5 mm thick laser-deposited layer composed of silicon, copper, and aluminium with the aim of increasing the alloy’s wear resistance. It is shown that the Al-Cu system is very sensitive to silicon additions, and that wear resistance depends on solidification of the primary phase to as well as on the final phase distribution. Development of a method for evaluating raw materials for use in iron ore sinter in terms of lime assimilation by W. Ferreira, R. Cromarty, and J. de Villiers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 325 A new test method is proposed that allows automatic evaluation of iron ores in terms of lime assimilation with increasing temperature. The new method, termed the Length Reducibility Test, is shown to be superior to the standard method in terms of precision and reproducibility of results, as well as ease of implementation. The recovery of manganese products from ferromanganese slag using a hydrometallurgical route by S.J. Baumgartner and D.R. Groot. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 331 The recovery of manganese metal and other manganese products from ferromanganese slag by means of leaching, precipitation, and electrowinning is investigated. The various methods are compared in terms of selectivity, costs, and product quality. Wear of magnesia-chrome refractory bricks as a function of matte temperature by M. Lange, A.M. Garbers-Craig, and R. Cromarty . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 341 The postulation that primary platinum group metal (PGM) matte will chemically react with magnesia-chrome bricks at temperatures above 1500°C was tested. Phase relations observed clearly indicate that chemical reactions take place between matte and the magnesia-chrome refractory under these conditions, and that these reactions are more complex than expected. Comparing the extent of the dissolution of copper-cobalt ores from the DRC Region using sulphuric acid in the presence of hydrogen peroxide and in tartaric acid by S. Stuurman, S. Ndlovu, and V. Sibanda . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 347 A copper-cobalt oxide ore from the Central African Copperbelt was leached in two different environments, and the effects of acid and reducing agent concentration and leaching temperature determined. The results indicate the potential of tartaric acid to extract cobalt, rather than copper, from copper-cobalt ores..
These papers will be available on the SAIMM website
http://www.saimm.co.za
A critical evaluation of haul truck tyre performance and management system at Rössing Uranium Mine by T.S. Kagogo* Paper written on project work carried out in partial fulfilment of B. Eng. (Mining Engineering)
The factors affecting haul truck tyre performance and the effectiveness of the management system at Rössing Uranium Mine were investigated with the aim of increasing tyre performance in terms of running hours until failure. The main objectives were to identify the types of tyre failure, their causes and cost implications, and evaluate the effectiveness of the management system. A site severity survey, weight study, and TKPH studies were conducted to determine the pit conditions, and an analysis of failed tyres carried out. The results showed that tyre performance at the mine has declined from 2009 to date, and the increase in lost value amounted to R5.7 million in 2012 alone. The main cause of tyre failure is loose rocks in the pit. The present management system in the load and haul department is not effective enough due to operational constraints it is facing. Keywords tyre failure, cost implications, types of failure, causes of failure, management system, TKPH, haul roads.
Overview Rössing Uranium Mine is an open pit mining operation on the west coast of Namibia. Currently, there is a global off-road tyre shortage, which has a negative impact on large operations such as Rössing. According to Cutler (2012), it has been estimated that most tyre suppliers have about 25–30% undersupply in the market for the past 3 years. The demand has surpassed supply, as shown in Figure 1, and the tyre shortage crisis has increased tyre prices by as much as 425% since 2009. Rössing currently has 32 Komatsu Haulpak 730E 2000 HP haul trucks which run on six tyres per truck, thus there are 192 tyres in operation at any given time. The downward trend in tyre life – 7371 hours in 2012 compared to 10 119 hours in 2009 (Figure 2) has a negative impact on operations at Rössing. Optimizing the tyre life is therefore a necessity in order to ensure that the operation does not run out of tyres, as obtaining tyres in the current market is not easy. According to Simulilo (2012), tyre performance at Rössing over the past two The Journal of The Southern African Institute of Mining and Metallurgy
Methodology and objectives The objectives and methodology of the project are summarized in Table I. The methodology for the project has been guided by the recommendations of Carter (2007) on areas of awareness for tyre improvement as shown in Table II.
Results and analysis Performance review The operation uses two main brands of tyres; namely Michelin and Bridgestone, but does use other brands such as Goodyear and Belshina when the supply of the two main brands is short. The operation is currently running 35 Komatsu 730E 180 t dump trucks, which use tyres sizes of 37.00R57 (Michelin) and 42/90R57 (Bridgestone). Figure 3 shows the frequency graph for the tyres in terms of the hours run to failure. Ideally the tyre life should fall in the
* Department of Mining Engineering, University of Pretoria, Pretoria, South Africa. © The Southern African Institute of Mining and Metallurgy, 2014. ISSN 2225-6253. Paper received Jan. 2014. VOLUME 114
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Synopsis
years has been below target, as illustrated in Figure 2. The mine has experienced premature failures in 49% of the tyres, with cut separation being the most common cause. The outcomes of the project will help the mine reduce the number of premature failures and increase the average tyre life. Rössing is currently under taking a costcutting exercise in the wake of poor production and low uranium prices coupled with increasing costs (Aluvilu, 2012). The project has the potential to assist in this, in line with the company’s objective, which is to cut costs without retrenching employees. Every additional hour of tyre life is money saved on the procurement of new tyres.
A critical evaluation of haul truck tyre performance and management system Table II
Areas of awareness for tyre life improvement (Carter, 2007) Awareness area
Possible actions to Improve awareness area
Driver awareness
• Make operators aware of the supply situation • Solicit input on areas of improvement • Provide incentives for improvements
Haul road design
• Super-elevation in corners (if supers aren’t possible, reduce speed) • Identify and remove soft spots in roads • Optimal road crown is 3%
Air pressure maintenance
• Conduct regular pressure checks, with immediate pressure corrections • Daily preferred, weekly necessity • Install new O-rings and hardware when mounting • Inspect/change/repair cracked wheels and components • Inquire with dealer about temperature/ pressure monitoring • Analyse air pressure documents just like any other
Mechanical maintenance
• Check alignment • Check suspension components • Use ‘rock knockers’ • Rectify problems immediately
Tyre and rim inspection
• Driver walk-around (train drivers what to look for) • Rim inspection for cracks or flange damage • Inspect valve hardware
TKPH management
• Total GVW adherence (no overloading is acceptable) • Adhere to speed limit
Support equipment
• Proper and effective use of graders and rubber-tired dozers • Equipment should be assigned to shovels • Driver radio communication of spills and road damage • Fix problem areas immediately
Analyse scrap tyres
• Analyse history of scrap tires • List types of damage • List vehicles with multiple tire failures • Examine shift performance (individual crews with problems; night vs. day)
Establish performance committee
• Involve cross-section of mine in joint efforts • Plan consistent meeting schedule • Make assignments for change; follow up for corrections
Communicate and report
• Issue consistent, visible reports of efforts • Issue consistent, visible reports of progress • Solicit suggestions.
Figure 1—Global tyre supply and demand (Cutler, 2012)
Figure 2—Tyre performance over the past 7 years
Table I
Objectives and methodologies Objective
Methodology
Conduct a market analysis on tyres.
Literature review
Do a performance review
• Data collection • Data analysis
Determine types of tyre failure at Rössing
• Data collection • Data analysis
Identify and investigate possible causes of tyre failure
• Road severity survey • Loading area investigation • Tipping area severity survey • Weight study • TKPH study • Interviews with the necessary personnel • Observation of general pit conditions
Determine cost implications of tyre failure
• Data collection • Data analysis
Evaluate the effectiveness of the Rössing tyre management system
• Operator questionnaire • Interview with necessary personnel • Data collection and analysis
Identify areas of awareness
Analysis and recommendations
10 000–12 000 hour interval to maximize tyre performance. However, in 2009 only 48% of tyres reached that interval, and only 18.7% and 2.5 % in 2011 and 2012, respectively. In 2011 and 2012 no tyres reached a life of more than 12 000 hours. In 2012 and 2011 most of the failures occurred between 6000 and 10 000 hours. 2009 has been chosen as a baseline year to compare 2011 and 2012 results because it is the best performing year over a 7-year period.
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Types of tyre failure Figure 4 summarizes the different types of tyre failure experienced at Rössing. Any of these modes of failure can occur at any time in the tyre’s lifespan except worn out which usually occurs at the back end of the service life of the tire. The biggest challenge the mine is facing is the fact that there is no system to identify specific areas in the pit where a particular tyre fails, hence the ‘hot spots’ where most tyre failures occur cannot be identified. This situation has made prioritizing specific areas in the pit for more attention and ascertaining the exact causes of tyre failure a hard task. The Journal of The Southern African Institute of Mining and Metallurgy
A critical evaluation of haul truck tyre performance and management system using the right tyres for the site conditions. TKPH causes tyre failure through heat separation and increased wear rate. The TKPH study was done only on the Michelin tyre and not the Bridgestone, but it was advised that the results for one brand are significant enough to determine the site conditions. Results (real site TKPH): ➤ Front= 706 ➤ Rear= 661.
Figure 3—Frequency graph for average tyre hours
The tyre in use is a 37.00R57 XDR B4 with a TKPH rating of 848, which means that the tyres on site are the right tyres; therefore the premature tyre failures are not due to the TKPH rating being exceeded. The results of the TKPH study are supported by the results of the weight study, which yielded the loading field data used in calculating the TKPH (TKPH = Average tyre load x Average truck speed).
Road corners
Figure 4—Failure modes
Most of the tyres failed from wear, which is the ideal situation as tyres usually have a lot of hours on them at the time of failure. In 2009 the fleet was a bit smaller than at present, but the fleet was the same for 2011 and 2012, which shows that more tyres were lost in 2012 than in 2011, indicating a drop in tyre performance as the operating truck hours were virtually the same for both years. Figure 5 shows the average tyre life for each failure mode for 2009, 2011, and 2012. The mine uses the operate-tofailure system whereby a tyre is replaced with a new tyre only once it has failed. This system does not pose any safety hazard as the radial tyres only deflate at failure and do not burst. Worn-out failure mode is the ideal type of tyre failure as it carries more hours and maximizes tyre life more than the other types of tyre failure. Operational failure and product-related failure remain the modes of failure with the lowest average hours across all three years, and much should be done to limit those modes of failure.
Road corners can be a source of tyre damage but this was not covered in the road severity survey, and an observation exercise was conducted to assess the conditions. Figure 6 illustrates loose rocks on the shoulder of a berm at a traffic circle and in a turn, which are a source of sidewall cuts. Figure 7 shows partially buried rocks in a turn, which are also a source of sidewall cuts and are not easily visible. The turning radius of the corners is within the mine standards, but the main concern is the loose rocks in the corners as the mine lost six tyres in 2012 due to sidewall cuts. The corners need to be constantly dressed with sand and loose rocks removed, and operators should take care not to cut corners and expose the tyres to rocks.
Loading and tipping areas Loading and tipping areas are among the areas that have the highest potential to cause tyre damage as they are a source of loose rocks that trucks can drive over. Rössing has two main
Possible causes of failure Driving over rocks is the general cause of most tyre failures at Rössing. There is no reliable system to locate where the failures take place in the pit, as tyres do not necessary fail immediately on coming into contact with the rock but might fail a few hours or days later. A few areas have been identified as causing specific modes of tyre failure, and the results from the observations and measurements are discussed in detail below.
Figure 5—Average hours per failure
Tons-kilometres per hour (TKPH)
The Journal of The Southern African Institute of Mining and Metallurgy
Figure 6—An example of loose rocks in a turn, with a truck about to cut a corner VOLUME 114
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A tons-kilometres per hour (TKPH) study was done at Rössing Uranium Mine to ascertain whether the mine is
A critical evaluation of haul truck tyre performance and management system Figure 10 shows an example of spillage on the in-pit haul road, which also is a major problem as it is a source of loose rocks that cause operational failures such as cut penetration.
Cost Implications The cost implication of premature tyre failure is based only on the remaining value in the planned tyre life, which constitutes a direct cost incurred. There are other added costs
Figure 7—Sidewall cut rocks in a turn
loading areas; namely the stockpiles and the blasted muckpiles in the pit. The two loading areas are serviced by shovels and front-end loaders with auxiliary equipment such as dozers for clean-up. Figure 8 shows a typical loading area condition at Rössing with loose rocks. Loose rocks are inevitable in the loading area, but should be cleaned up regularly and truck operators should not reverse into a loading area that contains loose rocks. Due to operational constraints, auxiliary clean-up equipment is not always available but effort should be put into keeping the area clean. The shovel operator should also practice a culture of cleaning the area regularly when the clean-up dozers are not available.
Haul roads A site severity survey was done to determine the pit conditions, including the haul road. The severity ratings are from 1 to 5, with 5 being the best condition and 1 the worst. Not all the categories are of significance to tyre failure, such as water and road width. The categories of high significance for tyre performance are spillage, gradient, banking, aggregates, undulations, and hammering, which all play a role in how the tyre fails and its wear rate. Spillage has a rating of 2, which means that around 50% of the haul road on average is covered with spillage and the rock size is in excess of 75 mm, which is large enough to cause tyre failure by modes such as rock cut penetration. Aggregates also have a rating of 2 and pose the same threat to tyres as spillage does. The gradient is rated 2 (about 10% both uphill and downhill), which can cause heat build-up in the tyre leading to failure through heat separation. Banking and hammering are of less concern as they have ratings of 3, which is good in the current state as tyres are subjected to shock less than 10% of the time and less than 50% of the roads have banking. Undulation is the biggest concern, with a rating of 1 due to the presence of undulations every 3 m. Undulations on the haul road causes payload spillages as well as a slight heat build-up in the tyre. Figure 9 shows an undulating portion of an in-pit haul road, clearly illustrating the severity of the situation.
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Figure 8—Loading area with loose rocks
Table III
Overview of in-pit haul road severity survey Category code
Category
Rating
Comment
CL493
Spillage
2
Spillage over 50% of the haul road, size in excess of 75 mm
CL494
Gradient
2
Gradient over 10% either uphill/downhill
CL495
Banking
3
Less than 50% of roads have banking
CL496
Aggregates
2
Aggregates cover 50% of the width of the haul road (over 25 mm)
CL497
Firmness
3
Sinkage is not in main working areas and is not affecting the tyres
CL498
Water
3
Standing water does not cause damage to the tyres
CL499
Undulations
1
Undulations every 3 m
CL500
Hammering
3
Tyres are subjected to shock damage less than 10% of the time
CL501
Road width
2
Equal to twice the width of a single truck
Figure 9—Undulation of in-pit haul roads at Rössing The Journal of The Southern African Institute of Mining and Metallurgy
A critical evaluation of haul truck tyre performance and management system department is facing. According to the superintendent of load and haul at Rössing (Fotolela, 2012) the department is facing constraints such as lack of operators and high production pressure, which leads to tyres getting minimum priority. Challenges facing the load and haul tyre management include:
Figure 10—Example of spillage on the in-pit haul road
incurred on top of the direct cost, such as the labour cost involved in changing tyres more frequently. A single tyre can take up to 8 hours to change, depending on which position it is in. A further indirect cost is the lost production time incurred by tyre changes. Figure 11 shows the remaining tyre value at failure for each type of tyre failure for 2011 and 2012. The highest cost was incurred due to rock cut penetration, which at R2 million is a 233.3% increase from R807 000 in 2011. Impact fracture is the other type of failure with increased lost tyre value from 2011 to 2012. Value lost due to failure by wear decreased from 2011 to 2012. No tyre value was lost in 2011 due to heat separation, while some value was lost due to heat separation in 2012. The highest costs are incurred through failure of tyres that have run the lowest number of hours which generally happens through premature failure. The cost per hour is obtained by dividing the value of a new tyre by the life in hours. In 2011 cut separation incurred the highest cost of R47.03 per hour, while failure due to wear had the lowest cost of R15.24 per hour. In 2012 impact fracture had the highest cost per hour with R47.50 per hour, while wear was the lowest with R18.27 per hour. The cost per hour results indicate that cut separation was the most expensive failure mode in 2011, while in 2012 impact fracture was the most expensive. The high cost per hour of premature failure is due to the low hours on the tyre at time of failure, while tyres that fail due to wear usually have had a long service, hence the low cost per hour. The average cost per hour for 2012 for all the failed tyres, irrespective of the failure mode, is R30.95 per hour, with the operational failures increasing the cost. The average cost is still well above the target cost of R24.00 per hour which is guided by the price of the new tyre and the target lifespan.
➤ ➤ ➤ ➤ ➤ ➤ ➤
Passive management No road maintenance programme No specific tyre preservation programme Low utilization of auxiliary equipment No system to identify point of failure in pit No record keeping of road maintenance Priority given to production at the expense of tyres.
According to Fotolela (2012) the department is working on addressing the challenges as tyre preservation has been identified as a high-priority area for 2013 in an effort to improve tyre performance.
Conclusion The mining industry has been facing a tyre shortage for the past 8 years, with a peak in 2008, and suppliers are currently running about 25–30% undersupply in the market for the past 3 years, hence the shortage is not expected to decrease until 2018. The mine experienced an upward trend in tyre life from 2006 to 2009 in terms of average hours per tyre, with 2009 having the highest average tyre life of 10 119 hours, while a downward trend has been evident since then with 7371 hours in 2012. In 2012 operational failure was the major failure category, accounting for 49% of all failures, and had the lowest average hours. The biggest concerns in terms of possible tyre failures are the conditions of the haul roads condition and loading areas, as they are the major sources of loose rocks that cause premature failure. The mine lost R5.7 million due to tyre failure in 2012, which is an increase from R4 million in 2011. The operational failure category made up the bulk of the tyre value lost, accounting for 73% of the total. The maintenance side of the tyre management system is very effective, as tyre maintenance practice, storage, and monitoring are within the Rio Tinto standards. The load and haul side of the tyre management system is not as effective, as no proper road maintenance record is kept, and effective utilization of clean-up dozers is low. This is due partly to the constraints, such as a shortage of operators, that the department is facing.
Load and haul tyre management system
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Figure 11—Lost tyre value for 2011–2012 VOLUME 114
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In the past, when the tyres were performing well, the load and haul management system had a bonus incentive for tyres, which engendered a positive attitude among the operators towards tyre preservation. Management also had more resources when it came to maintaining the roads, as Dust Aside was used more, which kept the haul roads in excellent condition. The load and haul tyre management is not highly effective due to different constraints that the
A critical evaluation of haul truck tyre performance and management system Recommendations 1. A tyre campaign should be adopted to promote collective responsibility and the spillage clean-up policy by all personnel reinstated. The campaign will not involve additional resources and can be very beneficial in terms of worker morale 2. A more aggressive approach should be taken to pit condition maintenance where resources permit and to overcoming the operational constraints that the auxiliary clean-up crews are facing in order to increase their effective utilization. This will be a challenge as the operation is facing an operator shortage, but efforts should be made to work with the available resources 3. A system is needed to identify the exact areas in the pit where tyre failure is common and classify them as ‘red zones’ that need priority in terms of auxiliary equipment.
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This will not be an easy system to come up with as most tyre failures do not immediately follow the cause. Aggressive road maintenance can help in this regard 4. The feasibility of re-introducing the tyre incentive programme should be investigated.
References ALUVILU, P. 2012. Production Engineer, Rössing Uranium Mine. Personal communication. CARTER, R. 2007. Maximizing mining tyre life. Engineering and Mining Journal, (00958948), vol. 208, no. 6. July/August. 58 p. CUTLER, T. 2012. EM tyre supply shortage – How it is affecting us’. AusIMM Technical Meeting, 14 May. FOTOLELA, D. 2013. Superintendent: Load & Haul, Rössing Uranium Mine. Personal communication SUMULILO, F. 2012. Maintenance Engineer, Rössing Uranium Mine. Personal communication. ◆
The Journal of The Southern African Institute of Mining and Metallurgy
A comparative study between shuttle cars and battery haulers by W.H. Holtzhausen* Paper written on project work carried out in partial fulfilment of B. Eng. (Mining Engineering)
with one using shuttle cars and one planning to change to shuttle cars. At 10 Shaft on the other hand, only shuttle cars are used.
Synopsis The purpose of this project was to compare two underground, batch coal haulers – battery haulers and shuttle cars – in order to determine the more viable machine to implement. The specific standards for battery haulers were investigated and compared to the requirements of shuttle cars in order to identify the unnecessary expenses related to the legal requirements that are attached to the machines. Costs such as running costs, capital costs, and maintenance costs were researched and compared over a typical life of machine. Average production rates and breakdown times were obtained and used to determine which machine would be more reliable in achieving the required annual production. Keywords coal, bord and pillar, hauling, cost (running, capital, total ownership), productivity, reliability, safety, availability, constraints.
Introduction to bord and pillar coal haulage In a typical underground mechanized bord and pillar mine section a Continuous Miner (CM) is used to cut the coal from the production face with a rooftogether bolting machine that installs permanent cemented roofbolts. Coal is loaded via the CM’s chain conveyor system onto the batch hauling machines, either shuttle cars or battery haulers, which then transport the coal load to an in-section crusher or feeder breaker where the coal is offloaded and crushed to a more tolerable size. From the feeder breaker, coal is transported out of the mine via a series of conveyor belt systems.
Background of the project
During studies at the University of Pretoria and field research at Exxaro’s Arnot coal mine, data was collected that could serve as a guideline for the selection of underground coal hauling machinery and what the present trends are related to these machines. Arnot Colliery is wholly owned by Exxaro Resources, the largest BEE (Black Economic Empowerment) contributing mining house in South Africa, and is situated in the Witbank Coal fields in the Highveld region. At present the mine is solely an underground operation, exploiting the No. 2 Lower seam with an average calorific value of 23.8 MJ/kg and ash content of 23%. Production is approximately 2.5 Mt/a and is supplied to Eskom’s Arnot power station on a cost-plus agreement (Exxaro, 2013). Arnot employs a mechanized bord and pillar method with continuous coal-cutting miners. At present there are two interlinked shafts, 8 Shaft servicing five sections and 10 Shaft four sections. Currently there are four sections at 8 Shaft that utilize battery haulers, The Journal of The Southern African Institute of Mining and Metallurgy
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Introduction and mine background
Batch hauling systems in an underground coal mine are the most unreliable link in the chain of ore transport. The implications of costs, productivity, reliability, and safety of these machines was scrutinized by the management of Arnot and it was found necessary to investigate these aspects further. During discussions, different opinions arose related to these machines and it was felt that these criteria need to be fully investigated and documented. In particular, a critical comparison between battery haulers and shuttle cars was required to determine the more feasible piece of equipment to implement in terms of cost, safety, reliability, and productivity.
A comparative study between shuttle cars and battery haulers Objectives and methodologies Objectives
Methodologies
Determine all factors that influence the advantages and disadvantages of battery haulers and shuttle cars
Collect information from different sources i.e. internet, management and operators. Use questionnaires as assistance
Determine safety and health issues Locate and consult articles about of both machines and evaluate them. the health and safety issues involved with battery haulers and shuttle. Determine what the industry trend is and what the future outlook is for the use of these machines
Contact manufacturers and request current and future sales profiles
Determine costs such as running, maintenance, capital and others that have a major influence on the use of battery haulers and shuttle cars
Acquire information from suppliers as well as mine employees on the costs involved with the haulers.
Obtain production figures of all the production sections for a certain time period and relate them to the batch haulers from that particular section
Enquire production figures from the surveying department of both 8 Shaft and 10 Shaft
Determine and compare the reliability and availability of the machines
Obtain downtime studies from the engineering department and calculate the relevant figures according to a specific standard
Evaluate and analyse the results
Study, in detail, all the results that were obtained and draw conclusions on findings. Calculate costs, production figures, availabilities and other relevant information
constant feed rate to the feeder breaker. Unlike battery haulers, there is no articulation, but steering is via all four wheels, allowing the vehicle to corner easier.
Results and analysis Current trends In the current South African coal mining industry Joy’s Stamler battery haulers are the front runners in the market and are used at most of the mines that use battery haulers. Joy is also the lead contracting company on the mine, supplying support and operations services to both the battery haulers and the shuttle cars. According to Stewart (2013) in the past 5 years Joy sold 15 battery haulers and 126 shuttle cars. During that time Joy had a backlog of three battery haulers and almost 100 shuttle cars. This is a clear indication that the industry is moving away from the use of battery haulers and is more prone to buying shuttle cars. The decrease in the use of battery haulers is resulting in a drop in the skills available for operation and maintenance, as well as increased cost of such skills. A key aspect to consider is the standardization of the mine fleet in order to ensure better focus and skills for a certain machine, effectively increasing operational life and performance.
Health and safety aspects Incident studies have shown that shuttle car operators have a reduced field of vision when the machine is loaded. In some cases operators lean out of the cab in order to see clearly, and
Battery haulers vs. shuttle cars The major difference between the two systems is that battery haulers are battery powered and shuttle cars are cable powered. This difference in itself entails advantages and disadvantages. Other factors are the coal load bearing methods, turning mechanisms, and flexibility in terms of use and transport. Battery haulers are much more flexible as they can travel any route in order to load and offload coal, but this can entail some drawbacks such as increased travelling distance and decreased battery life. This is not the case with shuttle cars since they are confined to travelling a specific route due to the trailing cable that supplies power to the drives. This, however, forces the section to do frequent section moves, where the section equipment is moved closer to the working face. Figure 1 depicts a typical battery hauler showing the articulation joint. This joint is very useful for cornering and manoeuverability in restricted conditions. The coal loading portion is confined to the rear of the machine and offloading is done with a hydraulic push-off system. This hydraulic mechanism is disadvantageous because it incurs spillages at the feeder breaker and therefore side plates have to be attached to the feeder. With the shuttle, car on the other hand, offloading can be synchronized with the feeder breaker since both use similar chain conveyor systems for loading and offloading (Figure 2). This reduces spillage and maintains a more
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Figure 1—Battery hauler (Joy Global, 2013)
Figure 2—Shuttle car (Joy Global, 2013) The Journal of The Southern African Institute of Mining and Metallurgy
A comparative study between shuttle cars and battery haulers this could result in injuries or fatalities. Another contributor to injuries is the cables that accompany the shuttle cars. These are tripping hazards, and when the haulers are in motion or are cornering these cables can come under tension, and when this tension is released injury, damage, or fatalities may also occur (Bezuidenhout, 2011) The major safety and environmental concern of the battery haulers is the fact that they are powered by lead-acid type batteries that produce hydrogen gases and might incur leakage of these gases and the acid-based electrolyte. This can lead to environmental contamination and exposure of workers to chemical-related injuries. The gases might also pose an explosion risk and this is of great concern in underground coal mines. (Van der Merwe, 2013)
Costs Costs are one of the most important factors in the present–mining environment day. Several cost categories have to be considered. These include procurement, operating or life–cycle costs, and the total ownership cost (TOC). Using the 2014 projected procurement prices that were supplied by Joy Global it was calculated that the difference in initial capital for the two machines was approximately 35%, with battery haulers being the more expensive. This excludes the additional requirement of installing battery bays as well as the ancillary ventilation and safety equipment required for the battery bay. The cost difference between the major overhauling and replacement bodies is 15% and 6% respectively, with battery haulers again being the more expensive. Considering that Joy requires these machines to receive a minor overhaul on a 1.0–1.2 million ton period, and a major overhaul every 1.6–1.8 million ROM tons, the life–cycle costs can be calculated using Figure 3. The average total life–cycle cost of the battery hauler per ton hauled is R 5.89 more than that of the shuttle car. With the projected annual requirement of 190 000 tons per month per machine, the average operating costs would be R3.89 per ton and R2.98 per ton for battery haulers and shuttle cars respectively. Figure 3 shows the life-cycle costs. The total machine life is approximately 2.4 Mt, which with the 190 000 ton per annum production target equates to approximately 12 years. Using this knowledge, a 10% interest rate, and a 2013/14 electricity cost of 65.51 cents per kilowatt-hour (Eskom, 2013) the TOC was calculated, making the assumption that these machines are operating for 18 hours a day and 26 days per month. Battery haulers have a total of 187 kW of motor power and shuttle cars 219 kW. Over this 12-year period the battery haulers will cost approximately R3.6 million more to operate moving the same volume of coal.
After seam height corrections had been applied, it seen that battery hauler sections produce on average 26 kt/month whereas shuttle car sections produce only an average of 21 kt/month. This is a difference of 5000 t/month per mining section. Also, battery haulers deviate from their monthly targets by about -20% and shuttle cars -30%, clearly indicating that there is a possibility that battery hauler sections are more productive.
Reliability Downtime data for all the machines was tabulated over a 1year period to obtain an average engineering availability for the two types of machines (Figure 4). An average of 96.7% and 96.2% availability is achieved for battery haulers and shuttle cars respectively with a standard deviation of 2% for shuttle cars and 1.3% for battery haulers, indicating that both machines are very close in terms of reliability. Battery haulers had an average of 191 hours per annum downtime and shuttle cars 165 hours. Figure 5 shows the specific downtimes for the haulers. ➤ Battery haulers: – Average mechanical downtime is 28.65% of the total – Electrical downtime is 54.9% – Hydraulic downtime totals 16.5%
Figure 3—Life-cycle costs
Productivity
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Figure 4—Engineering availability. Blue - battery haulers, grey - shuttle cars VOLUME 114
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The productivity of the different sections over a six-month period was considered to try to identify whether shuttle car or battery hauler sections have a higher productivity. Several other factors could also contribute to differences in productivity, such as the fact that battery haulers on average carry more tons.
A comparative study between shuttle cars and battery haulers Conclusions Table I is a representation, according to level of significance, of which machine has superior performance in terms of the various criteria.
Figure 5—Average downtimes of shuttle cars and battery haulers
➤ In terms of safety and environmental concerns the shuttle car is a much better option ➤ Shuttle cars incur fewer costs in both running and procuring ➤ Battery haulers offer better productivity ➤ Both machines are very reliable ➤ Shuttle cars are more adaptable and compatible with the other equipment in the transport chain ➤ Battery haulers are prone to be much more flexible ➤ The coal mining industry seems to shy away from the use of battery haulers.
Recommendations ➤ Shuttle cars: – Mechanical downtime is 48.1% of the total – Electrical downtime is 45.3% – Hydraulic downtime is 6.6%. Despite the belief that battery haulers are simpler in terms of the electrical circuitry, they experience higher relative downtimes due to electrical problems compared to shuttle cars. This may be due mainly to the Lionetics upgrades that are done on the machines to transform the DC current from the batteries into an AC output. This must be done according to SANS 1654. Also, battery haulers suffer more hydraulic breakdowns than shuttle cars and this raises concerns about the efficacy of the articulation joint. The relative mechanical downtimes of shuttle cars are higher, and this is because some of these machines are nearly 30 years old. However, they still show very high availabilities, similar to those of the battery haulers. Another factor that needs to be considered is that operators, artisans, and technicians have much more experience with shuttle cars than on battery haulers.
According to the results obtained from this study, it is recommended that the mine moves to the use of shuttle cars in order to reduce annual running costs, and also to reduce the funding required for battery bays, battery bay personnel, and refurbishing of batteries. In doing this the mine will also standardize the fleet and simplify ordering of spares and equipment. However, use of battery haulers should not necessarily be completely eliminated. In some cases, such as poor in-section and low seam conditions, battery hauler are the preferred choice because of the better flexibility and manoeverability. Thus further study into the effects that bad section and seam conditions have on the haulers is required.
Constraints Another factor contributing to machine selection is compatibility with other equipment such as feeder breakers and continuous miners. Through calculations and time studies it was found that the average cycle time for battery shuttle is 168 seconds and that of the shuttle cars is 190 seconds. Assuming that battery hauler payload is 18 t and the shuttle car payload 16 t with 21 cycles per hour for battery haulers and 18 cycles per hour for shuttle cars, that battery haulers have a 378 t/h capacity and shuttle cars only 288 t/h. With a maximum of three haulers in a section the maximum capacity in a section is 1134 t/h for battery haulers and 864 t/h for shuttle cars. Comparing these figures to the capacities of the feeder breaker (770 t/h) and the CM (840 t/h) shows that there will be some waiting time at the CM for both machines, but because of the increased flexibility of the battery haulers they will thus be underutilized. This clearly shows that shuttle cars are more than capable of transporting the required tons if they are maintained properly and that battery haulers will not be used to their full capacity. (Figure 6).
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Figure 6—Evaluation of capacities
Table I
Conclusive comparison Criteria
Weighting Shuttle car Product Battery hauler Product
Safety
10
8
80
6
60
Cost
9
8
72
6
54
Productivity
8
7
56
8
64
Reliabilty
7
9
63
9
63
Constraints
4
9
36
7
28
Flexibility
3
5
15
9
27
Trend
2
9
18
3
Total
340
6 302
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A comparative study between shuttle cars and battery haulers Acknowledgements ➤ My mentor and supervisor, Mr Nico van der Merwe, for his guidance and support throughout the course of the study ➤ The team at Exxaro Arnot for assistance with the research and the grant of this learning opportunity. ➤ The University of Pretoria’s Mining Engineering Department for their guidance ➤ Special thanks to Wilma De Jager at Arnot for her assistance in collecting and organizing of the data.
References ANGLO AMERICAN. 2013. Operations. www.angloamerican.co.za [Accessed 7 Aug. 2013]. BEZUIDENHOUT, L.J.A. 2011. Operational Report Mpumalanga Region February 2011. Department: Mineral Resources, Witbank. Eskom. 2013. Tariffs and Charges 2013/14. www.eskom.co.za [Accessed 11 Jul. 2013].
KENTUCKY COAL EDUCATION. 2013. Glossary of Mining Terms. www.coaleducation.org [Accessed 25 Feb. 2013]. KLINKERT, D and MARAIS, S. 2012. Equipment Replacement. Mobile Equipment Report. Arnot Coal. LEES, D. 2009. Slope instability. Exxaro Arnot. Code of Practice ARNOT_SMCOP_09/1. Arnot Reliability Centre. MTHETHWA, D. 2012. Battery Charging Bay Procedure. AR-BCB 01. Arnot Coal Mine. 12 December 2012. MYORS, A. 1998. Introduction of battery powered coal haulers into board and pillar panel production. Coal Operators’ Conference, University of Wollongong. O’DONNELL, C. and SCHIEMANN, M. 2008. Hydrogen gas management for flooded lead acid batteries. Battcon Stationary Battery Conference 2008. Mesa Technical Assosiates, Inc. SAMBO, V. 2013. Personal Communications. STEWART, D.N. 2013. Equipment Sales trend at Joy for 5 Years. Personal communication. VAN DER MERWE, N.J. 2013. Personal communications. NKOSI, S. 2007. South African Coal Industry – challenges and opportunities. Coaltrans Conference, Rome, 21-24 October 2007.
JOY GLOBAL. 2013. Haulage Systems. Product Overview. http://www.joy.com [Accessed 18 Jan. 2013].
UNIVERSAL COAL. 2012. Coal Mining In South Africa. http://www.universalcoal. com/projects/coal-mining-in-south-africa. [Accessed 8 Jul. 2013]. ◆
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EXXARO RESOURCES. 2013. Arnot http://www.exxaro.com/content/ ops/coal_arnot.asp [Accessed 17 Jan. 2013].
Investigation of cavity formation in lump coal in the context of underground coal gasification by C. Hsu*, P.T. Davies*, N.J. Wagner*, and S. Kauchali* Paper written on project work carried out in partial fulfilment of BSc. Eng (Mining)
Synopsis Underground coal gasification (UCG) is becoming more popular as the reserves of good quality, mineable coal are starting to diminish, and yet the global energy demand from coal is still increasing. The purpose of this research project was to investigate cavity formation within a coal block due to the combustion reactions in the context of UCG. The cavity plays a pivotal role in the UCG process, as it is essentially the gasification reactor. Cavity formation in an in situ gasification process using the forward combustion linking method (FCL) had been investigated, and a laboratory model was created to simulate the process. The experiment was performed by drilling a U-shaped tunnel into a coal block, which was then combusted internally with air that was fed through an injection hole. A heating element (at approximately 500°C) was used to supply the required heat for combustion at the base of the injection well. The coal blocks were analysed using micro-focus X-ray tomography. The tomography results showed that the coal tended to crack along the bedding plane after a short duration of combustion, due to either the formation of clinker or the expansion of swelling vitrinite along the horizontal tunnel. The deposit was thicker at the base of the injection well compared to the base of the production well; this may have been caused by the turbulence of the air flow and the relatively high oxygen concentration at the base of the injection well. A comparison of the results with work by Daggupati et al. (2010) showed the same trend, despite the slightly different methodology applied. Keywords underground coal gasification, cavity formation, gasification reactor, forward combustion linking method (FCL), clinker, vitrinite, micro-focus X-ray tomography.
UCG is an in situ gasification process. There are many methods for the gasification of coal seams and the extraction of syngas; the utilization of an injection and production well is investigated in this report. The injection well allows for the feed flow of air/oxygen/oxygenenriched air and steam, and the produced syngas is extracted to the surface via the production well. However, a linkage is required between the two wells as the coal seams are not sufficiently permeable to allow for the dispersion of air/steam across the horizontal length. There are several methods for creating this linkage path, namely: reverse combustion linking (RCL), forward combustion linking (FCL), hydro-fracking, electro-linking, and inseam linking (Abdul, Aqeel, and Gholamreza, 2013). The experiment was based on in situ gasification using the injection and production well with the FCL method. FCL involves the use of directional drilling, which offers the least damage to the existing ground structure (Directional Drill Pty Ltd, 2011) and improves the feasibility, design, and operation of a UCG plant. The purpose of this research was to investigate cavity formation in a coal block in the context of UCG. The work of Daggupati et al. (2010) was used as a guide for the experimental investigation and X-ray tomography was used to confirm and analyse the cavity formation.
Materials and methods Introduction
Experimental setup
Underground coal gasification (UCG) creates an opportunity to access coal that is deemed un-mineable due to its depth, the low quality of the coal, and unsafe mining conditions. In addition, it is considered to be more environmentally friendly than conventional mining and the process is less capital- and labourintensive (Self, Reddy, and Rosen, 2012). The produced syngas can be used by a host of technologies, but it is used mainly for the production of liquid fuels, power generation, industrial heating, or fertilizers (Seddon and Clarke, 2011).
The experimental setup was based on that of Daggupati et al. (2010), with the main
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* University of the Witwatersrand, Johannesburg, South Africa. © The Southern African Institute of Mining and Metallurgy, 2014. ISSN 2225-6253. Paper received Jan. 2014.
Investigation of cavity formation in lump coal in the context of underground coal gasification difference being the use of a heating element as the heat source in place of the LPG used by Daggupati et al. A Ushaped tunnel was drilled into a 15 cm × 11 cm × 12 cm coal block; the vertical injection and production wells were 5 mm in diameter and the horizontal tunnel connecting the two was 8 mm diameter. Stainless steel pipes were inserted into the vertical wells, one of which was connected to a gas rotameter and the air cylinder, and the other allowed the venting of the product gases into the fume hood. A Type K thermocouple was used inserted into the coal block to monitor the temperature throughout the experiment. Cutile refractory bricks were clamped together and surrounded the experiment in order to minimize heat loss and to keep the hightemperature zone separated from the external surroundings. The heat source was provided by a heating element made of 2 m of Kanthal D wire with 8.25 Ω/m resistance; the wire was wound into a spiral around a 3 mm rod. This wire was connected to a Variac (POWERSTAT® Variable Autotransformer) to control the voltage applied to the element and hence the temperature. The experimental setup is shown in Figure 1.
Experimental procedure An inlet air flow was set and maintained at 25 mL/s; this was fed into the injection well and monitored by a rotameter. The heating element was switched on using the Variac and the temperature was measured using a thermocouple. The experiment was allowed to run for 3 hours and the resulting coal block was analysed using the Nikon XTH 225 ST microfocus X-ray tomography system at South African Nuclear Energy Corporation (NECSA) (Hoffman and De Beer, 2013). This produced 3D images of the coal samples and allowed the mapping of the cavity based on density differences.
vitrinite by volume. It is important to note that all of the analyses were done using crushed coal samples, whereas the experimental work utilized large blocks of coal. However, the analysis can be assumed to be representative of the coal block as a whole.
Experiment 1 Expansion due to either the formation of clinker or the expansion of vitrinite caused a large crack to form along the length of the coal block during the first experiment. Clinker is a porous substance that forms during the combustion of coal that has a high ash content, non-combustible minerals, and a low gross calorific value (Mitra, 2011). The clinker is created by the fusion of non-combustible minerals (such as iron, calcium, and sodium) upon exposure to high temperatures (Afri Coal Investments Pty Ltd, 2013). Vitrinite is one of the organic substances found within coal and it is known to swell upon heating. Figures 2 and 3 show areas of darker grey which represent regions with a lower density material than coal. The low-density region is indicative of either clinker or plasticized vitrinite. For the purpose of this report, these regions will be referred to as ‘clinker’ as the material has not yet been analysed. More clinker formed around the inlet gas well (the left well) than the production well (right well); this is also shown in Figures 4 and Figure 5, where the shape of the clinker zone is shown. A higher concentration of oxygen was available at the inlet well, and the extent of the combustion reactions as well as the temperature was therefore higher
Table I
Coal characteristics Results and discussion
Component
A total of three experiments were performed, two of which yielded successful results, and which are referred to as Experiment 1 and Experiment 2.
Coal characterization The characteristics of the coal that was used are summarized in Table I. The ash content was18.8%, and from the maceral analysis it was found that the coal was composed of 80.8%
Figure 1—Diagrammatic representation of experimental setup (side view)
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Carbon Hydrogen Nitrogen Oxygen Sulphur Ash Moisture
Standard
Composition (%)
ASTM D 5373 ASTM D 5373 ASTM D 5373 ASTM D 5373 ASTM D 4239-05 ISO 1171:2010 ISO 11722:1999
71.05 4.29 1.78 1.69 0.99 18.8 1.4
Figure 2—Cross-sectional front view of combusted coal The Journal of The Southern African Institute of Mining and Metallurgy
Investigation of cavity formation in lump coal in the context of underground coal gasification Heat was provided to the coal for 30 minutes, but the air flow continued for a longer period. It was observed that the temperature continued to rise for a further 30 minutes until the self-heating system halted due to the eventual cooling of the coal surface by the air flow, which entered the coal at 23°C. When coal reaches temperatures of approximately 600°C, a self-sustaining heating will be established (Hoffman, 2008).
Experiment 2
Figure 3—Cross-sectional top view of combusted coal
The coal used for Experiment 2 was cemented into the refractory chamber, whereas the coal in Experiment 1 was only surrounded by refractory sand; this prevented the coal from cracking to such an extent. The shape of the clinker as shown in Figures 6 and 7 is similar to that seen in Experiment 1, however it is more cylindrical. Upon closer inspection (see Figures 8 and 9) it can be seen that the clinker layer is in fact slightly thicker at the base of the injection well (the well on the left of each photograph) than the production well (on the right).
Figure 4—3D cross-sectional model of combusted coal with clinker formation
Figure 6—Cutaway section of the coal showing the clinker formation
Figure 5—3D model of clinker formed in the coal after the duration of the experiment
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Figure 7—The shape of the clinker formation VOLUME 114
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than at the production well. The permeability of the clinker allowed oxygen to infiltrate deeper into and react with the coal (Wakatsuki, et al., 2009); this was notable particularly at the base of the injection well where the air flow was turbulent. The height difference between the clinker layers at each of the two wells is 8 mm and the width difference is 26 mm due to the large crack that formed where the surface area of coal is in contact with air (see Figures 2 and 3). The teardrop shape of the clinker gives an indication of the air flow pattern in the coal.
Investigation of cavity formation in lump coal in the context of underground coal gasification Conclusions and recommendations
Figure 8—Cross-sectional front view of combusted coal with dimensions
A laboratory-scale method was used successfully to simulate cavity formation during underground coal gasification. It was found that cracking tends to occur along the bedding plane of the coal due to the formation of clinker or expanded vitrinite, which forms in a thicker layer at the base of the injection well than at the base of the production well. This can be attributed to the turbulence of the air flow as well as the relatively high oxygen concentration at the base of the injection well. The air flow pattern has a great effect on cavity formation. The combustion products were not analysed, but should be in future work. It is recommended that further work be done on the heating source, such as using an alternative method. The use of Liquefied Petroleum Gas (LPG) is recommended as this method allows the experiment to mimic more of an industrial UCG process, in addition, if the temparature of the cavity could be controlled more effectively the clinker formation should be able to be minimized. Further investigations should take changes in the process variables into account; these variables include varying the injection air flow rate, the type of coal, duration of heating and the temperature of the heating source. Using steam together with the air would allow gasification reactions to occur in addition to combustion reactions, and thus mimic actual UCG processes more accurately.
Figure 9—Cross-sectional top view of combusted coal
Comparison to the research by Daggupati et al The current research was roughly based on the work done by Daggupati et al. (2010), the major differences between the two investigations being the type and size of the coal used, the heat source adopted, and the experimental duration. However, the general trends of cavity formation can still be compared. Figures 10 and 11 compare the top views of the cavity formation in the experiment by Daggupati and Experiment 1 respectively. The injection well is indicated by the white circle on the left of each image and the production well is on the right. It is clear that in both cases the cavity at the base of the injection well is larger than at the production well. This shape is not as obvious in Figure 11 due to the fact that a heating element was used across the entire length of the horizontal tunnel, causing clinker to form. It is also noteworthy that ash was formed in Daggupati’s experiment, whereas clinker was formed in the current investigation. This can be attributed to the different coal characteristics and operating conditions.
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Figure 10—Plan view of the cavity obtained by Daggupati et al. (2010)
Figure 11—Tomogram showing the plan view of the cavity produced in Experiment 1 The Journal of The Southern African Institute of Mining and Metallurgy
Investigation of cavity formation in lump coal in the context of underground coal gasification Acknowledgements The authors would like to extent their appreciation to their supervisor, Professor Wagner; to the workshop manager, Mr Samuel-McRae; to Doctor Kauchali and to Mr Hoffman and Mr de Beer at the micro-focus X-ray facility at Necsa.
References
http://www.directional-drill.com/horizontal-directional-drilling.html [Accessed 26 Oct. 2013]. HOFFMAN, G. 2008. Natural clinker. http://geoinfo.nmt.edu/staff/hoffman/ clinker.html [Accessed 18 Oct. 2013]. HOFFMAN, J.W. and DE BEER, F. 2013. Micro-focus X-ray tomography facility (MIXRAD) at NECSA. Personal communication. MITRA, S.K. 2011. Coal testing and analysis. http://www.mitrask.com/ coaltesting-analysis/index.html [Accessed 24 Oct. 2013]. SEDDON, D. and CLARKE, M. 2011. Underground coal gasification (UCG), its potential prospects and its challenges. http://www.duncanseddon.com/ underground-coal-gasification-ucg-potential-prospects-and-challenges [Accessed 25 Apr. 2013].
AFRI COAL INVESTMENTS PTY LTD. 2013. What causes clinkers in coal fired boilers. http://www.africoal.co.za/what-causes-clinkers-in-coal-fired-boilers [Accessed 18 Oct. 2013].
SELF, S.J., REDDY, B.V., and ROSEN, M.A. 2012. Review of underground coal gasification technologies and carbon capture. International Journal of Energy and Environmental Engineering, vol. 3, no. 16. pp. 1–8
DAGGUPATI, S., GANESH, A.,MANDAPATI, R.N., MAHAJANI, S.M., MATHUR, D.K., SHARMA, R.K., and AGHALAYAM, P. 2010. Laboratory studies on combustion cavity growth in lignite coal blocks in the context of underground coal gasification. Energy, vol. 35. pp. 2374–2386.
WAKATSUKI, Y., HYODO, M., YOSHIMOTO, N., NAKASHITA, A., and IKEDA, T. 2009. Material characteristics of clinker ash and examination of applicability for embankment. Powders and grains 2009: Proceedings of The 6th International Conference on Micromechanics of Granular Media, Golden, CO, 13–17 July 2009. http://scitation.aip.org/content/aip/proceeding/ aipcp/10.1063/1.3179872 [Accessed 24 Oct. 2013]. ◆
DIRECTIONAL DRILL PTY LTD. 2011. What is horizontal directional drilling.
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ABDUL, W.B., AQEEL A.B., and GHOLAMREZA, Z. 2013. Underground coal gasification: From fundamentals to applications. Progress in Energy and Combustion Science, vol. 39. pp. 197–199.
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Arnot’s readiness to prevent a Pike River disaster by R. Weber* Paper written on project work carried out in partial fulfilment of B. Eng. (Mining Engineering)
Methane explosions in underground coal mines are a major concern across the mining industry. After the Pike River disaster, Arnot Coal became even more aware of the explosion risk. To determine whether Arnot has adequate precautions against a methane explosion, such as the one that led to the Pike River disaster, a literature survey covering the causes of the incident and what preventative measures should have been in place was conducted. Many different methane explosion prevention methods were evaluated, as well as codes of practice. Based on the findings, Arnot’s measures to prevent conditions favourable for a methane explosion were evaluated. Keywords methane explosion, code of practice, dominoes, ventilation, LTR(last through road).
Project background What is coal seam methane? Methane (CH4) is formed as part of the process of coal formation. When coal is mined methane is eventually released from the freshly broken coal face. Methane can also be released as a result of natural erosion or faulting. The depth of the seam predicts the amount of methane content present. Methane is directly exposed to fresh air when mining takes place on surface and is confined when mining takes place underground..
Methane gas in coal mines Methane gas in underground coal mining is a big concern. Methane explosions are devastating, causing significant loss of life and damage to property, and there is a significant industry effort to prevent these accidents from occurring. Methane becomes explosive only if it is diluted to between 5%–15% by volume in air. Failure to provide adequate ventilation to The Journal of The Southern African Institute of Mining and Metallurgy
The Pike River disaster The Pike River disaster shocked the world. On 19 November 2010 at 3:45 pm there was an underground methane explosion at the Pike River coal mine which resulted in the loss of 29 lives. Daniel Rockhouse and Russel Smith were the only two people underground that survived the explosion. The emergency response was led by the New Zealand police. A rescue attempt was prevented by a lack of information regarding the conditions underground. On 24 November a second explosion occurred. and all hopes of finding the 29 miners underground alive were abandoned. The focus moved to the recovering of the bodies. However, conditions underground made this impossible. Two further explosions occurred, the second of which ignited the coal underground. The mine entrances were sealed in January 2011. This event raised concerns throughout the coal mining industry to prevent methane explosions. (Royal Commission of the Pike River Coal Mine Tragedydegy, October 2012)
Scope of study The aim of the study at Arnot was the prevention of methane explosions and not preventing a methane explosion from leading to a coal dust explosion. Thus the project’s main emphasis was on what happened at Pike River, how the tragedy could have been prevented; what measures Arnot already has
* Department of Mining Engineering, University of Pretoria. © The Southern African Institute of Mining and Metallurgy, 2014. ISSN 2225-6253. Paper received Jan. 2014. VOLUME 114
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Synopsis
dilute the methane to less than 5% by volume increases the threat of an explosion. Following the Pike River disaster in New Zealand in November 2010, it became a major concern at Arnot 10 Shaft to manage its methane levels so as to avoid a similar incident.
Arnot’s readiness to prevent a Pike River disaster Abandoned panel ventilation
Table I
Depth vs. methane content (Van der Merwe, 2013) Depth interval (metres)
Mean methane content (cubic metres per ton of coal)
100 500 1000 1500 2000
0.02 0.99 3.73 4.89 7.09
➤ COP requirement—Abandoned panels should be sealed off or ventilated until they are sealed off. Where they are still ventilated, the air flow velocity in the LTR should be greater than 0.5 m/s ➤ Actual—From Figure 4 it can be seen that panel N31 is abandoned and thus sealed off as required by the COP. Figure 5 shows an abandoned panel that is still being ventilated prior to sealing off. The air flow velocity is 1.1 m/s, thus complying with COP.
in place, and what further measures it needs to take in order to prevent a similar disaster. There was no emphasis on actions to be taken in the event of a methane explosion, only on prevention. The project also did not consider rescue procedures or assistance to the bereaved families.
Methodology A survey was conducted and a hierarchy of information on methane explosions built up by using methods such as consulting with experts in the field and contacting other mines that had suffered methane explosions. From the hierarchy the most relevant information was selected. Underground visits were conducted during December 2012 and June 2013. The visits were scheduled so as to ensure that enough time was spend at each installation to ensure the effectiveness of Arnot’s methane management and explosion preventative measures. Following the evaluation, conclusions were drawn and the necessary measures put into place to make Arnot ‘Pike -River ready’.
Figure 1—Mining layout showing barrier pillars
Results and analysis Ventilation requirements ➤ COP requirement—Barrier pillars should be spaced 15 m apart. ➤ Actual—From Figure 1 it can be seen that the barrier pillar spacing separating panels averages 23.35 m, thus complying with the COP ➤ COP requirement—Each section should have its own ventilation district with a separate set of ventilation controls ➤ Actual—It can be seen from Figure 2 that Section 4 and Section 12 have separate ventilation districts, which means each section has separate intakes as well as return airways. The air flows are kept separate by making use of walls and air crossings, thus complying with the COP ➤ COP requirement—Under normal conditions 0.25 m3/s of fresh air should be supplied in the last through-road, and 0.3 m3/s under high-risk conditions ➤ Actual—Taking the average bord height and width as 3 m and 6.5 m respectively, the following air flows were calculated: Normal conditions = 7 faces * 3 m * 6.5 m * 0.25 m3/s = 43.2 m3/s High- risk conditions = 7 faces * 3 m * 6.5 m * 0.3 m3/s = 41 m3/s From Figure 3 it can be seen that that sections 4 and 12 comply with the COP. The air flow in Section 3 was measured as 46.9 m3/s, thus also complying with the COP.
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Figure 2—Air-crossing separating ventilation districts
Figure 3—Section 4 and Section 12 layout and air flow rates The Journal of The Southern African Institute of Mining and Metallurgy
Arnot’s readiness to prevent a Pike River disaster Schedule of checks Table IV shows the inspection intervals prescribed by the COP, together with the actual inspection intervals noted during underground visits. It must be noted that the inspection intervals are those prescribed under normal conditions, although there could be exceptions.
Analysis of results ‘Swiss cheese’ model of causation Each layer in a ‘Swiss cheese model’ (Figure 6) represents a defensive system labelled by type (at the top). The holes in each layer represent gaps in the defensive system. These gaps can be created by active failures, human error, violations etc. Once these gaps line up there is no defence and an accident such as the Pike River disaster is likely to occur. For example, if the ventilation at Pike River had been adequate and if there had been no ignition sources, then the accident would not have occurred. However, both the ventilation and the spark prevention measures were inadequate. The more defensive systems in place, the better the chances that not all of the holes in the model will line up (the probability of an incident decreases). It is thus of great importance to have as many defensive systems as possible. If the critical systems fail, the secondary or ancillary systems must kick.
Figure 4—Example of sealed and unsealed abandoned panels
Dominoes at Pike vs. Arnot mandatory Code of Practice
Figure 5—Air velocity in last through-road of abandoned panel
From Table V it can be seen that 24 dominoes were identified that led to the Pike River disaster. Visual inspections, calculations, and interviews with employees indicated that Arnot is able to prevent all 24 of these dominoes.
Intakes and returns roads minimum velocities
Conclusion
Table II and Table III show the average intake and return velocities for each section. All of the readings can be seen to comply with the COP requirement of 0.5 m/s.
Failure to control the methane levels resulted in led to the inevitable explosion at Pike River during 19 November 2010, which resulted in the unnecessary loss of 29 lives. By
Table II
Intake velocities Section 3 4 12
COP intake requirement (m/s)
Section minimum measurement (m/s)
Average intake velocity (m/s)
0.5 0.5 0.5
1 1 1
1.3 1.2 1.1
COP return requirement (m/s)
Section minimum measurement (m/s)
Average return velocity (m/s)
0.5 0.5 0.5
1.2 1.3 1.3
1.7 1.9 1.7
Table III
Section 3 4 12
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Return velocities
Arnot’s readiness to prevent a Pike River disaster Table IV
Inspection intervals Inspected
According to COP
Inspections interval measured
Compliance with COP?
Main ventilation to section
Start of shift and then every three hours
Start of shift and then every three hours. Deviations were less than 10 minutes.
✓
Last through-road velocity
Start of shift and then every three hours
Start of shift and then every three
✓
hours. Deviations were less than 10 minutes. Positive ventilation of faces
Start of shift and then every three hours
Start of shift and then every three hours. Deviations were less than 10 minutes.
✓
Scrubber screen and fan on the
Start of shift and when changing picks
Start of shift and every time picks
✓
mechanical miner
are changed Start of shift
Start of shift
✓
Trailing cables
Start of shift and at least once during shift
Start of shift and mostly once, sometimes twice.
✓
Flame proofing
Start of shift visual inspection
Start of shift
✓
Operating conditions of spray nozzles
Start of shift and when changing picks
Start of shift and mostly once, sometimes twice.
✓
Mechanical miner onboard flammable gas monitor tested
Start of shift
Start of shift
✓
Test for flammable gasin each heading up to second-last row of support
Start of shift and then every three hours
Ventilation brattices and section walls installed according to standard
Start of shift and then every three hours. Deviations were less than 10 minutes
✓
Figure 6—Swiss cheese model (Bredel, 2013)
thoroughly working through the Pike River disaster, 24 main dominoes contributing to the Pike River disaster wincident as were identifiedlisted. It can be seen that ventilation practisces was the main issue involved. Following up on a review of Arnot’s methane control systems, which were obtained and data of which included obtaining from the ventilation readings, conducting visual underground inspections, and exercising to communicate with the interviews with relevant persons at the mine itself, Arnot’s practices were compared to those of Pike River. From these comparisons it can be seen that Arnot can prevent all of the twenty four 24 main dominoes that played a role at Pike River. The fact that Pike River was 150 m deep, compared with only 60 m at Arnot, could also have played a role. The mean
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methane content measured in cubic metres per ton of coal increases with increasing depth of the mine. The methane emission rate would therefore have been higher at Pike River than at Arnot. (This can be seen from Table I).
Recommendations Increase the scoop/line brattice efficiency The scoop/line brattice efficiency can be calculated by comparing the quantity of air entering the section (point 1, Figure 7) to the quantity of air leaving the section (point 2). As an example, if the amount of air entering is 1.6 m/s and the amount leaving is 1.6 m/s, the efficiency will amount to 100 per cent. The Journal of The Southern African Institute of Mining and Metallurgy
Arnot’s readiness to prevent a Pike River disaster Table VI
Efficiency table Section
Above 70%
Below 70%
Average%
36 8 24
18 46 30
75 64 68
3 4 12
It is recommended that the efficiency be maintained equal to or greater than 70 per cent. From Table VI it can be recommended that the scoop/line brattice efficiency on Section 4 must be drastically increased, since this low efficiency results in less air being supplied in the last through-road. Section 12 can also consider improving their efficiency. Section 3 has the highest efficiency, but it is still nevertheless recommended that more effort be put into minimizing the number of readings below 70 per cent efficiency.
Figure 7—Layout of a basic section at Arnot
Table V
Dominoes Reporting format Domino
Complies with COP
It is recommended that I.O (in order) and O.O.O (out of order) should not be used to report on ventilation readings in the shift overseer’s daily logbook. The quantity of the readings should rather be recorded so as to build up a record of ventilation readings underground.
Pike River
Arnot
Proper ventilation of goaf area
x
✓
Sealing of the hydro panel
x
✓
Unrealistic drive towards coal production
x
✓
Main fan placed underground
x
✓
Suggestions for further work
In-bye ventilation fragile
x
✓
Ventilation engineer
x
✓
Second intake
x
✓
Assessing of Health and Safety reports
x
✓
Free venting of drained methane
x
✓
Tripping of machines
x
✓
Sensor bypassing
x
✓
Experience of mine management/
x
As we have learned, by placing the main fan at the PR in the underground vicinity was ‘a major error’. It appears that the safety measures similar in Australia (but not legal requirements in NZ) were not enforced nor instituted to begin with. It is thus suggested that the aforementioned and vital legislation in different countries regarding the prevention of methane explosions be investigated. In addition, investigations should be extended to the prevention of coal dust explosions, and not only methane explosions. As with main fans that are not banned underground within the NewZealand laws it is suggest that there could be further looked into the role the laws of the different countries play when it comes to the prevention of methane explosions. It can also be suggested that it must be further looked at the prevention of a coal dust explosion and not only that of a methane explosion
✓
constant change Sensor at bottom of vent shaft
x
✓
Placement of the sensors and reporting
x
✓
Sufficient sensors
x
✓
Voltage cables near pipes transporting
x
✓
x
✓
x
✓
methane creating a hazard Stoppings separating
References
intake air from return air
down of air while work continuous
x
✓
Variable-speed drives
x
✓
Training of employees
x
✓
Widening of the panel from 30m to 45m
x
✓
High methane readings reported
x x
✓
Risk assessment conducted when fixing restricted and nonrestricted areas.
Main fan flameproof
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✓
BREDEL, P. 2013. Risk [Interview] (14 Apr. 2013). Senior lecturer, University of Pretoria MINING-TECHNOLOGY. 2012. mining-technology. http://www.mining technology.com/projects/pikeriverminenewzeal/pikeriverminenewzeal3. html [Accessed 23 Jan. 2013]. Royal Commision of the Pike River Coal Mine Tradegy, October 2012. Report Volume 1, Wellington, New-Zealand. VAN DER MERWE, N. 2013. Mineral recource manager, Exxaro, Arnot. WIKIPEDIA. 2013. Wikipedia. http://en.wikipedia.org/wiki/Pike_River_Mine [Accessed 23 Jan. 2013]. VOLUME 114
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Diverting air away from face and shutting
Organised by
On behalf of
Primary sponsor
Younger o Members
Committee
INNOV VAT TE RES SPECT COMPETE
Supported by
Laser cladding AA2014 with a Al-Cu-Si compound for increased wear resistance by K.J. Kruger* and M. du Toit* Paper written on project work carried out in partial fulfilment of B. Eng. (Metallurgical Engineering)
Aluminium alloys have gained popularity in many industries due to their high strength and low weight. One shortcoming of aluminium alloys is their poor resistance to abrasion and erosion wear compared to materials such as stainless steels. In this project, aluminium alloy 2014 (AA2014) was coated with a 1.5 mm thick laser-deposited layer composed of silicon, copper, and aluminium with the aim of increasing the wear resistance. The amount of silicon, copper, and aluminium added to each sample was determined by a mixtures model. It was discovered that the Al-Cu system is very sensitive to silicon additions and that wear resistance depends on the primary phase to solidify as well as on the final phase distribution. Two primary phases were identified; alpha aluminium and theta intermetallic. It was observed that the clad layer increases both the hardness and wear resistance of AA2014, and that the material solidifying as primary alpha aluminium displayed a lower hardness but higher wear resistance than the samples containing primary theta phase. All clad layers performed better in terms of wear resistance than the unclad samples. The knowledge gained and principles used in this project could be applied to many other aluminium alloys. Keywords aluminium alloys, laser cladding, wear resistance.
Introduction AA2014 is a lightweight and high-strength material that has proven invaluable in applications that require a high strength-toweight ratio. The alloy finds most of its applications in the aerospace industry, although new inroads into the drill piping industry have been established. Due to the low alloying element content in AA2014, the material displays poor hardness and, therefore, poor wear resistance. There is therefore a need to increase the wear resistance of AA2014 without altering the bulk mechanical properties such as ductility and strength-to-weight ratio.
Scientific background Due to the low alloying element content, AA2014 is not particularly hard and is therefore not abrasion resistant. This is the case for most 2000 series aluminium alloys, The Journal of The Southern African Institute of Mining and Metallurgy
* Department of Materials Science and Metallurgical Engineering, University of Pretoria, Pretoria, South Africa. Š The Southern African Institute of Mining and Metallurgy, 2014. ISSN 2225-6253. Paper received Jan. 2014. VOLUME 114
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Synopsis
and it has become desirable for industrial application to increase the wear resistance of these alloys without compromising the material’s strength or strength-to-weight ratio. The alloy composition and relevant mechanical properties are shown in Tables I and II respectively. A potential solution to the low wear resistance of AA2014 is to coat the material with a wear-resistant layer. This project will consider the use of laser cladding an Al-Cu-Si compound onto the AA2014 substrate with the intention of increasing the wear resistance of the surface while leaving the bulk mechanical properties of the material unchanged. The aim of this cladding process is to produce a harder material by intermetallic phase formation between aluminium and copper in the form of Al2Cu. This intermetallic phase formation as well as the microstructure can be seen in Figure 1. This microstructure was achieved using a 40 wt% Cu 60 wt% Al clad composition (Dubourg, 2002). It was found that the intermetallic phase was the first to solidify, and the resulting hardness and abrasion resistance was the highest in the alloy range that was tested. In the current investigation, silicon was added in an attempt to minimize porosity in the cladding. If successful, this cladding process could be applied to many aluminium alloys in which wear resistance is important but impractical to achieve by conventional methods. Tough alloys that do not necessarily display good wear characteristics could be surfaced using this process to increase the resistance to abrasive wear while maintaining the toughness of the bulk material.
Laser cladding AA2014 with an Al-Cu-Si compound for increased wear resistance Table I
Typical chemical composition (in weight %) of aluminium alloy AA2014. Single values denote maximum limits. (Capalex, 2013) %Cu 3.8–4.9
%Fe
%Si
%Mn
%Mg
%Cr
%Zn
%Ti
%Other
%Al
0.5
0.5
0.3–0.9
1.2–1.8
0.1
0.25
0.15
0.15
Balance
Table II
Typical mechanical properties of AA2014 in the T4, T351 heat-treated condition (Kaufman, 2002) Ultimate tensile strength (MPa) 472
Yield strength (MPa)
Brinell hardness number (500 kg/ 10 mm)
Modulus of elasticity (GPa)
325
120
73
Figure 1—Microstructure of laser-clad 40%Cu-Al material forming primary intermetallics and a secondary eutectic phase etched with Keller’s reagent (Dubourg, 2002)
the case of laser cladding, the focal point of the beam is slightly above the work piece so as to melt the feed material without melting much of the substrate. This causes minimal dilution and rapid solidification of the feed material onto the substrate while ensuring that a metallurgical bond is achieved. These properties are desirable for the abovementioned application. This will result in little to no change in the substrate bulk properties while achieving the desired result of depositing a layer of feed material that will act as an abrasive-resistant coating. A powder feeder was used to feed the material into the weld pool via a carrier gas. The powder feeder method was chosen over other methods, such as preplaced powders, because it is the only method that is industrially viable (Vilar, 2001). Figure 2 illustrates the process of laser cladding using a powder feed. It can be observed from the schematic that the powder is fed into the weld pool via a powder injection nozzle and is transported in space by a carrier gas. Once the powder enters the weld pool, it melts and will rapidly solidify to form a coating as the laser moves along the track.
Cladding material An important microstructural aspect to consider when selecting materials suitable for wear resistance is the presence of a finely distributed hard phase in a matrix of a more ductile phase. It is also important to ensure that the
Literature review Laser cladding Laser cladding is a low heat input, rapid solidification process. The advantages of this process include high levels of productivity due to fast welding speeds, low levels of dilution, refined microstructures due to rapid solidification, minimal distortion of the work piece, and small heat-affected zones (HAZs) (CSIR, 2012). There is a general consensus on the advantages of laser cladding in these respects, regardless of the cladding feed material or substrate being clad (Liu, 1995; Hyatt, 1998; CSIR, 2012; Joining Technologies, 2012). This process makes it possible to bond a material with a completely different microstructure and mechanical properties to a substrate. The process makes use of high-intensity light emission that is focused using several lenses to a point above, below, or on the surface of the work piece depending on the intended application (Joining Technologies, 2012). In
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Figure 2—Schematic of laser cladding using a powder feeder (Vilar, 2001) The Journal of The Southern African Institute of Mining and Metallurgy
Laser cladding AA2014 with an Al-Cu-Si compound for increased wear resistance cladding material is compatible with the substrate and that no brittle phases are formed along the fusion line that could result in poor metallurgical bonding. When considering an abrasion-resistant layer, the most important aspect is the cladding hardness. There is a general consensus in the literature that material hardness is directly proportional to abrasion resistance (Kim, 2006; Jeong, 2003; Caldwell, 1988). This is due to the fact that harder materials are more resistant to plastic deformation and will therefore resist the process of wear more effectively than softer material. Figure 3 illustrates the effect of hardness on the wear resistance of pure metals and several steel alloys. From Figure 3 it is clear that, in general, as the hardness of a material increases, so does its abrasion wear resistance. It can be observed from the results of the study by Dubourg (2002) that the wear resistance of the layer is directly related to cladding hardness. Figure 4 is a graphical representation of the wear resistance of an aluminium sample as the copper content is increased. Figure 5 shows a graphical representation of the hardness of the clad material measured at varying depths prior to any post-cladding heat treatment (Dubourg, 2002). It is evident that an increase in copper content increases the hardness of the clad layer. The material displays a rapid drop in hardness at a depth of approximately 0.9 mm, most likely due to the transition into the virgin aluminium base material. Figure 5 indicates that the hardness of the cladding increases with copper content. This is due to an increase in the amount and particle size of the intermetallic θ phase. Figure 6 displays the equilibrium Al-Cu binary phase diagram. It can be seen that the eutectic composition is at 33 wt% Cu; however, this true only for the binary system under equilibrium conditions. Pseudo-binary Al-Cu-Si phase diagrams are considered in Figures 7 and 8 at 1% and 10% silicon respectively. It is clear that as the silicon concentration of a sample is increased, the
formation of primary theta phase is favoured at lower concentrations of copper. The formation of either primary alpha aluminium or primary theta phase could affect the mechanical properties of the material and must therefore be carefully monitored.
Figure 4—Graphical representation of the decrease in wear rate under different wear parameters with increasing copper content in the cladding material (Dubourg, 2002)
Figure 5—Graphical representation of hardness for various cladding compositions as a function of depth (Dubourg, 2002)
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Figure 6—Aluminium-copper phase diagram (Murray, 1992) VOLUME 114
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Figure 3—The relationship between wear resistance and hardness in several pure metals and alloys (Tylczak, 1992)
Laser cladding AA2014 with an Al-Cu-Si compound for increased wear resistance using optical emission spectrometry (OES), as optical microscopy, and scanning electron microscopy usin energydispersive spectroscopy (SEM-EDS). Based on the results of these tests, two samples were selected for wear testing, together with the unclad control sample.
Results and discussion The results from the OES chemical analysis, micro-Vickers hardness tests, and SEM-EDS phase analysis are displayed in Table V.
Chemical analysis Figure 7—Vertical section of the Al-Cu-Si phase diagram containing 1% Si (Raghavan, 2007)
It is evident from the Table V that while the silicon content was acceptable in each case, the copper content in all samples was consistently lower than the designed copper content. This implies a problem with the copper delivery system. A possible solution would be pre-mixing of the powders and using only one hopper. The system is very sensitive to silicon content and for this reason, unexpected solidification patterns were observed. Several samples solidified as primary α-Al while others solidified as primary θ-intermetallic. Localized EDS measurements indicated that neither the primary nor the eutectic phase contained silicon values higher than 1%. Microstructural examination of the cladding was therefore carried out.
Microstructure The addition of silicon promotes the formation of primary θ
Figure 8—Vertical section of the Al-Cu-Si phase diagram containing 10% Si (Raghavan, 2007)
Methodology
Example of a table used to capture experimental data (weight %) (Balance is aluminium) Sample no.
Twelve samples were laser-clad with layers containing from 30–54 wt% copper and 0–3.4 wt% silicon (Table III). One sample remained unclad and was subjected to the same testing procedure as a control: The layer compositions were selected using a statistical method known as the mixture lattice design. This design was chosen to maximize the useful data that is extractable from the experiments and allow for the mapping of material characteristics based on chemistries using contour lines.. The centre point in the experiment (samples 1, 6, and 12) was performed in triplicate in order to test the consistency of the cladding process. The samples were produced by gas-feeding three powders (pure Cu, pure Al, and 12%Si-Al alloy) from three hoppers into the weld pool via a triaxial nozzle. The laser parameters used to produce the samples are listed in Table IV. These parameters produced a smooth, well-bonded clad layer, but have not been optimized for all the samples. After polishing, some unmelted particles were observed in the inter-bead region of most samples. The clad samples were subjected to chemical analysis, micro-Vickers hardness tests, and microstructural analysis
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Si content in clad (%)
Cu content in clad (%)
0 1.1 2.3 3.4 0 2.3 4.5 6.8 0 1.1 2.3 3.4
Unclad control 30 37.1 44.3 51.5 42 44.3 46.5 48.8 54 49.2 44.3 39.5
1 2 3 4 5 6 7 8 9 10 11 12 13
Table IV
Final laser parameters Laser parameter
Final value
Power Spot size Step-over distance Resultant overlap Welding speed Gas flow rate (1.5 l/min per hopper)
2.5 kW 3 mm 0.8 mm 73.3% 1.5 m/min 4.5 l/min
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Laser cladding AA2014 with an Al-Cu-Si compound for increased wear resistance Table V
Results of chemical analysis, hardness tests, and phase analysis Sample no.
1 2 3 4 5 6 7 8 9 10 11 12 13
Design Si content in clad (wt %)
Design Cu content in clad (wt %)
Hardness (HV)
OES analysed Si content (wt %)
OES analysed Cu content (wt %)
Phases present
2.3 1.1 0 3.4 0 2.3 4.5 6.8 0 1.1 3.4 2.3 Control
44.3 37.1 30 51.5 42 44.3 46.5 48.8 54 49.2 39.5 44.3 Control
243 234 214 370 277 225 259 250 260 265 203 199 116
2.2 1.1 0.2 3.1 0.2 2.1 4.2 6 0.2 1.2 3.2 2.1
33.3 28.5 22.3 39.7 31.6 31.8 35.4 33.5 41.5 39.2 32 33
Al+θ Al+θ Al+θ Al+θ+Si Al+θ Al+θ Al+θ+Si Al+θ+Si Al+θ Al+θ Al+θ+Si Al+θ
phase, which was apparent in the microstructure of most samples (as seen in Figure 9). However, the primary α phase formed in several samples, (Figure 10). It can be seen from Figure 9 that there is a large amount of primary θ phase and very small amounts of eutectic phase in the microstructure of sample 9. This is considered to be the desired microstructure due to the large amounts of hard intermetallic phase surrounded by a more ductile eutectic phase. As shown in Figure 10, sample 12 displays primary phase solidification and this results in soft particles that are surrounded by the harder (α-Al +θ) eutectic phase.
Samples containing higher amounts of silicon (3% and more) contained a third phase, consisting of small grey particles that were visible under an optical microscope. An example of these particles can be seen in Figure 11. These small grey particles were difficult to identify using SEM back-scatter electron imaging (BEI), and a chemical map was constructed in order to identify the locations of the particles. The results can be seen in Figure 12.
Figure 11—Sample 4, containing tertiary grey phase Figure 9—Sample 9, displaying primary theta solidification
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Figure 12—Chemical map of sample 8 VOLUME 114
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Figure 10—Sample 12 displaying primary alpha solidification
Laser cladding AA2014 with an Al-Cu-Si compound for increased wear resistance It is clear that there are areas of high silicon concentration in this sample and these areas cannot be distinguished from the rest of the microstructure by SEM. The particles could not be accurately analysed using EDS due to their small size, and thermodynamic modelling using FactSage™ was used to predict the equilibrium phases that form on solidification in the samples. The third phase that forms in the high-silicon samples was identified as high-purity (>99%) silicon particles. When the fusion line was inspected, it was evident that very good metallurgical bonding between the clad layer and the metal substrate had been achieved. This is evident in Figure 13. The formation of grain boundary precipitates can also be observed; however, these precipitates could be eliminated through solution annealing.
sample 4 had a higher hardness than sample 12, it had a higher mass loss during the wear test. This can probably be attributed to spalling at the sample edges (as shown in Figure 15).
Mechanical properties Hardness is believed to have the largest influence on the wear resistance of the material, and thus all the samples were subjected to fifteen hardness measurements using the microVickers hardness test. The results were collated and used to generate a hardness contour map with relation to individual sample chemistries, using ‘Design Expert’ (Figure 14). It can be seen that high harness is due to a synergistic effect between copper and silicon.
Figure 14—Hardness (HV) contour map based on chemistries obtained from OES results
Wear testing
Table VI
Based on the results obtained from the above tests, samples 12 and 4 as well as the unclad control sample were subjected to wear testing. Sample 12, which solidified as primary α-Al, displayed the lowest hardness, and sample 4, which solidified as primary θ-intermetallic, displayed the highest hardness. A slurry erosion wear test proved to be too uncontrollable to provide consistent results, and therefore a non-standard, wet two-body abrasion wear test was performed. Tables VI and VII describe the test conditions and results. It is important to note that the wear test was extremely aggressive and high mass loss was observed. Sample 12 displayed the highest wear resistance, while the unclad sample displayed the lowest wear resistance. Although
Test conditions in non-standard wet two-body abrasion wear test Parameter
Value
Abrasive medium Force applied to sample Run time Rotation speed
Diamond-impregnated resin (220 grit) 100 N 10 min 310 r/min
Table VII
Mass loss in two-body abrasion wear test Sample no.
Mass loss per unit area exposed (g/cm2)
Control sample (unclad) Sample 4 (hardest) Sample 12 (softest)
Figure 13—Fusion line between cladding layer (top) and substrate (bottom)
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0.162 0.156 0.134
Figure 15—Spalling of sample 4 The Journal of The Southern African Institute of Mining and Metallurgy
Laser cladding AA2014 with an Al-Cu-Si compound for increased wear resistance Conclusions
CAPALEX. 2013. 2014 alloy data sheet. http://www.capalex.co.uk/alloy_types/ 2014_alloy.html [Accessed 15 May 2013].
The surface hardness of AA2014 was successfully increased by laser cladding. The effects of the cladding operation were examined using a variety of tests, and an understanding of the mechanisms by which the cladding layer will protect the base material from an abrasive environment has been gained. This cladding process was successful in providing additional wear resistance as well as hardness to the substrate material without reducing the strength-to-weight ratio of the material. The wear resistance of the material showed a strong correlation to the primary phase that solidifies, and the solidification mechanism is linked to both the thermodynamics and kinetics. It was determined that primary alpha aluminium solidifying samples are more wearresistant, but have lower hardness than the primary theta phase intermetallic solidifying samples. The reason for the low wear resistance of the hard primary theta solidifying samples is most likely due to spalling. Furthermore, it was determined that both the primary intermetallic and primary alpha aluminium samples were harder and more wearresistant than the substrate. This implies that the process could be applied to a wide range of aluminium alloys and will broaden the range of application of aluminium alloys by increasing the lifespan of these materials under severe wear conditions.
CSIR. 2012. Laser welding. http://www.csir.co.za/lasers/laser_welding.html [Accessed 13 May 2013]. DUBOURG, L., PELLETIER, H., VAISSIERE, D., HLAWKA, F., and CORNET, A. 2002. Mechanical characterisation of laser surface alloyed aluminium–copper systems. Wear, vol. 253, no. 9. pp. 1077–1085. HYATT, C.V. 1998. The Effect of Heat Input on the Microstructure and Properties of Nickel Aluminum Bronze Laser Clad with a Consumable of Composition Cu-9.0Al-4.6Ni-3.9Fe-1.2Mn. A Metallurgical and Materials Transactions. JEONG, D.H. 2003. The relationship between hardness and abrasive wear resistance of electrodeposited nanocrystalline Ni–P coatings. Scripta Materialia. pp. 1067–1072. JOINING TECHNOLOGIES. 2012. Laser beam welding. http://www.joiningtech.com/ industry-references/welding-types/laser-beam-welding [Accessed 9 May 2013]. KAUFMAN, J.G. 2002. Aluminum alloys. Handbook of Materials Selection. Kutz, M. (ed.). John Wiley & Sons. Ch. 4. p. 104. KIM, K.T. 2006. KYUNG TAE KIM, S. I. 2006. Hardness and wear resistance of carbon nanotube reinforced Cu matrix nanocomposites. Materials Science and Engineering. pp. 46–50. LIU, Y. 1995. Microstructural Study of the Interface in Laser-Clad Ni-AI Bronze on AI Alloy AA333 and Its Relation to Cracking. A Metallurgical and Materials Transactions. Murray, J. 1992. Binary alloy phase diagrams. Introduction to Alloy Phase Diagrams. Vol. 3. Baker, H. (ed.). ASM Handbook, ASM International, Materials Park, OH. p. 2.44. RAGHAVAN, V. 2007. Al-Cu-Si. Journal of Phase Equalibria and Diffusion. pp. 180–181. TYLCZAK, J.H. 1992. Abrasive wear. Friction, Lubrication, and Wear Technology. ASM Handbook, vol. 18. Materials Park, OH. pp. 184–190.
CALDWELL, S.G. 1988. A microscopic study of the behavior of selected Al-Cu alloys in unlubricated sliding wear*. Wear. pp. 225–249.
VILAR, R. 1999. Laser cladding. Journal of Laser Applications, vol. 11, no. 64. pp. 64-81. ◆
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References
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Development of a method for evaluating raw materials for use in iron ore sinter in terms of lime assimilation by W. Ferreira*, R. Cromarty*, and J. de Villiers* Paper written on project work carried out in partial fulfilment of B. Eng (Metallurgy)
Steel is produced in a basic oxygen furnace from hot metal obtained from a blast furnace. A sintered iron ore with good hightemperature properties (strength and permeability) should be used as feed to the blast furnace. The quality of this sintered ore depends on the reactivity of the iron ore used as feed to the sinter plant during the lime assimilation step in the sintering process. The penetration test is the standard method for evaluating the reactivity of iron ore with lime. It is, however, difficult to determine the exact depth of penetration from the standard test. A new test method is proposed that allows automatic evaluation of iron ores in terms of lime assimilation with increasing temperature. A comparison of the coefficients of variation for the new and standard methods for each ore type demonstrates that the results of the new test are more reproducible and more precise than those of the standard method. The test is also less time-consuming and easier to implement. Keywords iron ore, sintering, properties, lime assimilation
Introductions It is important to determine whether iron ore used as feed to a sinter plant will produce a sinter suitable for use in a blast furnace. The feed to the blast furnace should be permeable (slightly porous, but not too porous since this could have an adverse effect on sinter properties) and have a high strength. A few standard tests are available to evaluate raw materials for use in the sinter process. The current method used can be difficult to interpret, time-consuming, and provides mostly limited information. A new test method, which allows calculation of the reactivity of iron ore with lime with increasing temperatures, was investigated. The new method should be easy to implement, reproducible, quick, and give valid results.
Background Kumba Iron Ore identified a need to develop a new test method to evaluate iron ore in terms of lime assimilation. Lime is added to iron ore to improve the porosity and to obtain the The Journal of The Southern African Institute of Mining and Metallurgy
Steelmaking Steel is produced by three processes, of which the blast furnace (BF) and basic oxygen furnace (BOF) combination has been most popular since the nineteenth century. Liquid pig iron (hot metal) produced by the BF is used as feed to the BOF. The solids feed to the BF include iron oxide, metallurgical coke, sintered ore, manganese ore, and dolomite (Buschow, 2001, pp. 4293–4296).
* Department of Materials Scnience and Metallurgical Engineering, University of Pretoria, Pretoria, South Africa. Š The Southern African Institute of Mining and Metallurgy, 2014. ISSN 2225-6253. Paper received Jan. 2014. VOLUME 114
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Synopsis
correct sinter strength and permeability for use in the blast furnace. If the sintered ore does not have adequate strength, a large amount of fines will be produced during stockpiling and handling of ore before use in the blast furnace. These fines will then be blown out by the offgas from the top of the blast furnace. Insufficient porosity of the sintered ore, on the other hand, will not allow sufficient flow of gas through the ore, causing insufficient reduction of iron in the blast furnace and resulting in a product that is not suitable for steelmaking. The porosity of the sintered ore should, however, not be too high since this can have an adverse effect on the sinter properties. Higuchi et al. (2003, p. 1388) found that different iron ore types react differently during lime assimilation, depending on surface morphology as well as chemical composition. The standard method for determining whether an ore type is suitable for use in the sinter plant is the penetration test. A 5 mm lime tablet is placed on top of a 10 mm iron ore tablet, both pressed from powder. The ore and lime tablets are heated in a furnace and the lime penetrates into the ore tablet. The depth of penetration is used as an indication of the reactivity of the ore type.
Development of a method for evaluating raw materials for use in iron ore sinter Blast furnace Solid material is fed into the BF from the top while hot gas is blown from the bottom upwards, as can be seen in Figure 1. This gas, which consists of a mixture of air and pure oxygen, is used to improve combustion efficiency to reduce the iron ore with metallurgical coke. Molten metal with a slag layer on top collects at the bottom of the BF and the liquid metal is tapped to be used as feed in the BOF. In order to obtain sufficient contact between the gas and solid particles in the BF, a permeable burden is required to allow a high and uniform gas flow rate (Barker et al (2006, p.1393). The iron feed material should not contain excess fines, since the fines will be lost to the top gas. Barker et al (2006, p.1393) also mention that sinter strength is an important characteristic since the sinter used as feed to the BF will be subjected to stockpiling, handling, and transportation. During all of these steps the sinter should not degrade and produce fines, which will be blown out of the BF with the top gas. To ensure that the iron ore feed has sufficient permeability, strength, and correct size, the fine ore is sintered.
Iron ore sintering Iron ore sintering can be described as the controlled burning of a fuel mixed with iron ore (Barker et al, 2006, p. 1393). The process converts natural fine iron ore material, screened iron ore fines, coke, and lime into a fused clinker-like aggregate that can be effectively used in the BF. Iron ore fines are mixed with 5% anthracite, which acts as fuel and is conveyed in the sinter through the process. The mixture of sinter material is fed onto a moving grate, and at the feed end gas burners are used to ignite the top of the bed. As the mixture moves forward, the combustion zone progresses downward due to air flow through the permeable bed. This results in a temperature profile through the thickness of the bed. Temperatures as high as 1300°C to 1480°C are reached at the hottest spot in the bed, causing the particles to fuse together into porous clinker. At the discharge end, sintering
would have occurred through the thickness of the bed. The material is then crushed and screened. The oversize material is sent to the BF stockpiling yard and the undersize is returned to the sinter process. Sintering of the ore has further benefits to the BF operation since the flux is incorporated into the sinter mix instead of being added separately to the BF feed, and it also produces a sized sinter as feedstock with better hightemperature properties (Barker et al, 2006, p. 1394).
Lime assimilation during iron ore sintering With the world-wide increasing demand for iron ore, more lower-quality grades of iron ore are being produced than in the past. As sinter plants have to make use of the ore at hand, it is important to have a quick pre-production test to evaluate whether an ore will be suitable for use in the BF after sintering. The iron ore sintering process consists of three stages. The first stage is the heating of the burden before melt formation takes place. The second stage is the primary melt formation of pseudo-particles with an adhering fines layer, and the last stage is the assimilation of lime with the nucleus ore (Hida and Nosaka, 2007, p. 103). Various tests are used to determine the properties of the ore during each step. In this project the focus will be on testing the properties obtained during the last stage of the sintering process. Three tests currently are used in industry to evaluate iron ore in terms of lime assimilation. These include the small packed bed sintering test and the penetration test as described by Higuchi et al (2004, pp. 1385-1386), and also the variation of the small packed bed sintering test (Hida and Nosaka, 2007, p. 104). Since the penetration test makes use of a tablet of iron ore and a second tablet of a combination of lime and iron ore, the conditions in this method are similar to those found in the sintering process. This method was therefore chosen as the standard to be used to evaluate the new developed test method, and is hence the only method described here in more detail.
The standard penetration test
Figure 1—Blast furnace feed materials (Buschow et al. 2001, p. 4296)
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The standard penetration test is used to evaluate the melting of fines into the adhering layer of pseudo-particles during sintering. The test uses two pressed tablets: an ore tablet consisting of an equivalent mass ratio of size fractions -0.25 mm and +0.25–0.5 mm and a primary melt tablet consisting of 26 mass% CaO-FeO). The ore of the two size fractions is mixed and pressed into a tablet 10 mm in diameter and 5 mm in height, while the primary melt reagents are hand-mixed for 20 minutes and pressed into a tablet with diameter and height of 5 mm. Both tablets are produced using an iron mould at a pressure of 0.314 kN. The primary melt tablet is placed on top of the ore tablet in the centre of a nickel vessel. The sample is heated from ambient temperature to 800°C within 3 minutes, then from 800°C to 1300°C in 2.5 minutes. The temperature is maintained at 1300°C for 2 minutes, after which the furnace is then set to cool over 10 minutes. A vertical section of the sample is mounted and polished. Macro images at 5× magnification are used to measure the depth of penetration, which is defined as the distance from The Journal of The Southern African Institute of Mining and Metallurgy
Development of a method for evaluating raw materials for use in iron ore sinter
New length reducibility test method Industry requires an easy test method to determine the assimilation of iron ore with lime during sintering, as there is a need to determine the reactivity of a given type of iron ore with lime and the subsequent porosity and sinter strength. Since the grade of iron ore being mined is decreasing due to the increase in demand, it is becoming even more important to establish whether sintering of a particular ore will result in a suitable product for use in the blast furnace. A disadvantage of the penetration test is that there is no certainty on the measured depth of penetration, since penetration invariably does not proceed in the perfectly hemispherical fashion described in the literature. The depth often varies across the section and there is uncertainty whether to measure the deepest, the shallowest, or average depth of penetration. Furthermore, the samples sometimes react to such an extent that a completely molten mass is obtained, and no conclusion can be drawn from such results. This test is also not considered to be fully reproducible, since measurement of the penetration depth is done manually. The alternative test method, which is aimed at achieving conditions similar to those found in the sintering process, makes use of a single cylindrical sample consisting of a mixture of 26 mass% lime and 74 mass% iron ore which is placed in the furnace. Hewakandamby et al. (2013, p. 456) described a test method to determine the behaviour of ash from biomass and coal at elevated temperatures. A cylindrical sample is heated in an ash fusion furnace, and images of the sample obtained at fixed temperature intervals are used to determine the temperatures at certain states, for example the softening temperature. In the new length reproducibility test method used here, the furnace is equipped with a digital camera to obtain images of the sample with increasing temperature. MatlabTM is then used to convert the digital images to a greyscale image to find a matrix equivalent to the size of each sample. As the samples are heated and started to react, the height of the samples decreases and can be measured as a function of temperature by the software program and a plot of height fraction versus temperature obtained. The samples are large enough to allow most samples to be submitted subsequently for additional testing such as sinter strength tests or optical examination to determine the porosity.
Figure 2—Vertical section through lime and iron ore pellet after sintering (Hida and Nosaka, 2007, p. 104) The Journal of The Southern African Institute of Mining and Metallurgy
Hewakandamby et al (2013, p. 454) concluded that the normal ash fusion test used in the past does not produce consistent results as it is based on visual inspection and the results are therefore dependent on the subjective judgement of the person carrying out the experiment. The original ash fusion test as applied to iron ore sintering cannot, therefore, be regarded as inherently reproducible. The method described here is an automated one, with a higher precision than that of the original visually evaluated ash fusion test.
Experimental procedure The validity of the new length reducibility test method was evaluated by comparing the results with those from a standard penetration test, using various ore types. Three ore types were tested: two haematite ores, a typical Northern Cape ore in South Africa, a West-African ore, and a goethitic iron ore from Marra Mamba in Australia. All three ore types were tested under the same conditions using both methods, including the quantities of ore and lime used to produce the samples. Eight samples of each ore type were tested in each of the two methods.
Standard test procedure The standard penetration test procedure used was based on the conditions given by Higuchi et al. (2004, pp. 13851386). The test was carried out in an infrared furnace, which has a fast enough heating rate to simulate the sinter temperature profile.
Samples and sample preparation All sampling methods and preparations were done according to ISO 10836 (ASTM part 12), which is equivalent to ASTM E877 (ASTM part 12). The two tablets were placed in the furnace simultaneously, one on top of the other. The top tablet was 5 mm in diameter and height and comprised 26 mass% lime and 74 mass% Fe ore. The bottom tablet, 10 mm in diameter and 5 mm in height, consisted of equal amounts of -0.25 mm and 0.25–0.5 mm pure iron ore. Given the density of iron ore fines of 5 g/cm3 (Fe ore MSDS), the total mass of iron ore needed for a 10 mm ore tablet was 1.96 g. In order to limit the number of variables in the test work, all the samples of each ore type were prepared from a single batch mixture. The total mass of iron ore required for the samples for each ore type was milled to the required particle size distribution (equal amounts of -0.5 +0.25 mm and -0.25 mm) and compressed at 0.314 kN in a hand-operated hydraulic press. The melt tablet consisted of 0.113 g lime and 0.322 g iron ore (5x5 mm site consisting out of 26% lime and 74% iron ore, The ore and lime for all of the samples for each ore type was hand-mixed in a single batch and pressed into a green pellet. When the correct particle size distribution of equal amounts of -0.5 mm +0.25 mm and -0.25 mm, was obtained, the mixed product was pressed into pellets in a handoperated hydraulic press at a compressive load of 0.314 kN. The samples were heated in an infrared furnace for the standard penetration test. The ore sample was placed in the furnace with the 5 mm melt tablet on top of the 10 mm ore tablet. The furnace was set to simulate the sinter process. The temperature was increased to 800°C over 3 minutes, then to VOLUME 114
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the top rim of the tablet to the tip of the reaction zone (Higuchi et al, 2004, p. 1386). Figure 2 shows the heated sample with the penetration depth indicated by the arrows. This measurement is used to compare the reactivity of different ore types (Higuchi and Okazaki, 2004, p. 432).
Development of a method for evaluating raw materials for use in iron ore sinter 1300°C in over 2.5 minutes, and maintained at 1300°C for 2 minutes. The temperature was then decreased to ambient over 10 minutes and the sample removed from the furnace. The reacted samples were sectioned vertically through the axial centre and mounted in resin. The bottom half of the sample was removed to allow inspection of the centre of the sample by conventional and stereo-optical microscopy. Since the 5× magnification of the sample surface did not allow full inspection of the surface, the stereo microscope was used to take three successive images of the entire surface area, which were combined into a single collage. A distinct defect or grain was used as reference point to allow the entire penetration depth to be measured from the bottom of the sample (ore tablet) up to the point of maximum penetration from the top. This total was then subtracted from the original height of the ore tablet to find the depth of penetration, which could be identified as a darker uniform phase at the top of the tablet, as seen in Figure 3. The bottom granular part consisted of iron ore particles with size fractions of equal amounts -0.25 mm and -0.5 + 0.25 mm.
Length reducibility test procedure These tests were carried out in an ash fusion furnace equipped with a digital camera to capture the change in height of the various samples as a function of temperature.
Samples and sample preparation The 10 mm diameter by 5 mm height tablets for these tests consisted of single samples with 26 mass% CaO and 74 mass% iron ore, i.e. a lime to ore ratio equivalent to the penetration tests. The ore comprised equal amounts of -0.25 mm and +0.25 -0.5 mm material, as was the case for the penetration test. The procedure for preparing the powders, mixing them, and pressing them into green pellets was identical to that used in the penetration test.
Once the furnace reached 1204°C it was switched off and allowed to cool to 400°C before the samples were removed. The temperature of 1204°C was chosen to ensure that the samples did not melt completely and molten material leak onto the furnace tube, thereby cracking it.
Analyses of the samples The digital images of the samples in the furnace were analysed using a program written in ImageJ specifically for this project. The program requires a manual but accurate rectangular selection of each sample in the image to be drawn based on the initial width and height of the sample. The selection allows the program to set boundaries on where these measurements need to be taken, which allows it to focus exactly on where the sample to be measured is located in the bigger image. The pixels in this rectangle will be counted. After this initial set-up, the image acquired at the maximum temperature is opened together with the image immediately preceding. The program then automatically determines the difference between the two images, thresholds the image, converts it to a binary image, and removes any outliers. These steps are illustrated in Figures 4 and 5. It can be seen that in the temperature increment from 1202°C to 1204°C, only the Northern Cape ore tablet showed a measurable change in area (Figure 5). After this step the program makes use of a matrix to measure the number of pixels that make up the difference in surface area measured (red pixels in Figure 5). This gives an indication of the total reactivity of the ore type during that temperature increment. The amount of pixels counted can now be divided by the initial width measured by the program from the rectangular
Heating procedure Up to four samples were placed simultaneously in the ash fusion furnace by balancing them on the sample carrier in the furnace, with the camera focused on all four samples. The furnace was set to increase the temperature at 7°C/minute and the computer set to take an image of the four samples with each 2°C increase in temperature, starting at 1150°C.
Figure 3—Marra Mamba pellet with darker melt phase penetrating into granular ore phase
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Figure 4—Typical Northern Cape ore (top left on stand) and WestAfrican ore (top right on stand ) iron ore at 1202°C (A) and 1204°C (B) in the ash fusion furnace
Figure 5—Difference between images A and B in Figure 4 (indicated by red) as measured by the macro written in Image J The Journal of The Southern African Institute of Mining and Metallurgy
Development of a method for evaluating raw materials for use in iron ore sinter selection to determine the height change during the temperature step. The data is displayed in a text file and can be imported directly into Excel® to construct a graph and to quantitatively determine the results. The greater the change in height for the sample, the more reactive the ore type. Since the reactivity analysis is done by making use of the images as well as the written program, the samples removed from the furnace can be subjected either to a strength test or a porosity measurement, as no further dimensional measurements need to be taken.
Additional results Figure 7 indicates that the Marra Mamba and Northern Cape ores start reacting at about 1185°C, while the West-African ore starts reacting only above 1200°C. This information cannot be obtained from the standard test.
Advantages and disadvantages In order to determine which test method would produce results faster and easier, the disadvantages and advantages of both methods are listed.
Results and discussion Eight samples of each of the three different ore types were tested by the two test methods. The average penetration depth measured for the penetration tests and average change in height for the length reducibility test were used to quantitatively compare and analyse the two test methods. The results are given in Table I
Standard test method and new test results
Standard test Advantages: ➤ Since this test is carried out in the infrared furnace, a high heating rate can be used, thus simulating the temperature profile of the sintering process more closely ➤ The infrared furnace cools down, and the sample can be removed, within an hour
Table I shows that the trends of the results obtained for both test methods were similar. Both methods indicated that the West-African ore had the lowest reactivity, followed by the Northern Cape ore with medium reactivity, and then Marra Mamba with the highest reactivity. These results show that the new length reducibility test can therefore be used to obtain the same type of information as the standard penetration test used in industry. Northern Cape ore
Reproducibility of test methods To determine the reproducibility of both methods, the coefficient of variation was plotted for each ore type. The coefficient of variation, Cv, is calculated by:
West-African ore
Marra Mamba
Figure 6—Coefficient of variation for each ore type tested using both methods
[1] where σ is the standard deviation and μ is the mean. A lower coefficient of variation indicates a higher reproducibility and more precise results. Figure 6 shows that the new length reducibility test has a consistently lower coefficient of variation than the standard penetration test, indicating that the new method has a higher reproducibility than the standard test, potentially leading to more precise results.
Typical Northern Cape ore West-African ore Marra Mamba
Figure 7—Average change in height as a function of temperature for the three ore types tested as determined by the length
Table I
Comparison of results
1 2 3 4 5 6 7 8 Average Standard deviation σ/μ
1.44 0.23 1.41 0.56 1.82 1.52 0.79 0.20 1.00 0.63 0.63
Total Penetration (Old Method) (mm)
West-African ore Marra Mamba Typical Northern Cape ore West-African ore 0.04 0.22 0.16 0.05 0.02 0.15 0.19 0.02 0.11 0.08 0.77
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1.55 0.74 1.21 1.27 0.83 1.02 0.86 1.07 1.07 0.27 0.25
0.36 0.64 0.62 1.38 0.48 0.48 1.21 2.69 0.98 0.78 0.80
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Marra Mamba 0.10 0.59 1.82 0.49 1.24 3.64 1.87 4.49 1.78 1.56 0.88
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Total change in height (New Method) (mm) Typical Northern Cape ore
Development of a method for evaluating raw materials for use in iron ore sinter ➤ Since the sample has to be mounted and sectioned in order to determine the reactivity, a porosity measurement can easily be carried out. Disadvantages: ➤ Only one sample can be tested at a time, making this a time-consuming method. ➤ All samples are mounted and vertically sectioned, and thus are not available for strength tests ➤ Working with the mould to produce the melt tablet requires great care since the 5 mm mould is extremely delicate and can resist hardly any pressure ➤ When placing the tablets in the tube of the infrared furnace, balancing the small melt tablet on top of the larger ore tablet is very difficult ➤ The depth of penetration is hard to measure. For some ore types a definite penetration can be seen as a uniform and darker phase penetrating into the granular phase of the ore, but this is not necessarily the case for all ore types ➤ The depth of penetration is measured by hand and the final value is often subjective.
tablet which is simple to reproduce. There is no need to produce two tablets, each with a different mixture of raw materials 2. Results are obtained automatically by making use of the images taken with the digital camera and the image analysis program. The results are therefore obtained easily and the error associated with human judgement is avoided.
Future work and recommendations In order to obtain international accreditation for this new method, the following aspects are of particular importance according to ISO 17025: ➤ ➤ ➤ ➤ ➤ ➤
Reproducibility/repeatability Discrimination of samples Sensitivity Detection threshold Comparison against existing methods Inter-laboratory tests.
Conclusion New length reducibility test Advantages: ➤ The ash fusion furnace allows testing of up to four samples per test run. The length reducibility test is therefore a more productive test ➤ The analysis of the reactivity is done automatically by a computer program and this removes operator subjectivity ➤ No vertical cross-section of the sample after the test is necessary, since the reactivity is a function of the change in height of the sample. The samples are thus available for either a strength test or a porosity measurement ➤ The sample consists of a single tablet that is easy to produce to the required weight and dimension specifications ➤ This method does not require a 5 mm diameter tablet made in a fragile mould. Disadvantages: ➤ The maximum temperature for these three ores could not be set higher than 1204°C, since at 1206°C the Marra Mamba ore melted completely. The melt could leak into the furnace tube, causing it to crack ➤ The ash fusion furnace has a very slow heating rate and cannot replicate the actual heating rate in the sinter process.
Apart from its proven greater reproducibility and precision, the length reducibility test also provides the following advantages: 1. The time required for sample preparation is reduced, since only one tablet is required instead of two for each ore type and test run. Four samples can be analysed simultaneously during a single 4-hour run, including cooling time, compared with two hours for a single sample sample by the standard method. The new test method also requires only one cylindrical
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References ASTM Standards, Chemical Analysis of Metals; Sampling and analysis of metal bearing ores. 1982. Part 12 E877. ASTM, Philadelphia. BARKER, J., KOGEL, J., KRUKOWSKI, S., and TRIVEDI, N. 2006. Industrial Minerals and Rocks - Commodities, Markets, and Uses. 7th edn. Society for Mining, Metallurgy, and Exploration (SME), Warrendale, PA.. BUSCHOW, K., CAHN, R., FLEMINGS, M., ILSCHNER, B., KRAMER, E., and MAHAJAN, S. 2001. Encyclopedia of Materials - Science and Technology, vols. 1-11. Elsevier, Amsterdam. FE ORE MSD UNIVERSAL MINERALS. HEWAKANDAMBY, B., LESTER, E., PANG, C.H., and WU, T. 2012. An automated ash fusion test for characterisation of the behavior of ashes from biomass and coal at elevated temperatures. Fuel, vol. 103. pp. 454–466. HIDA, Y. and NOSAKA, N. 2007. Evaluation of iron ore fines from the viewpoint
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In a comparison of the standard test with the newly proven length reducibility test it was found that: 1. The new length reducibility test yields more reproducible results 2. The results obtained by the new test method are more precise 3. The new length reducibility test is less timeconsuming and easy to implement 4. The new test can therefore be used with confidence to evaluate and quantify the reactivity of iron ore in terms of lime assimilation in the sintering process.
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of their metallurgical properties in the sintering process. Transactions of the Institute of Mining and Metallurgy C, Mineral Processing and Extractive Metallurgy, vol.116. pp. 101–107. HIGUCHI, K., HOSOTANI, Y., OKAZAKI, J., and SHINAGAWA, K. 2003. Influence of iron ore characteristics on penetrating behavior of melt into ore layer. ISIJ International, vol. 43. pp. 1384–1392. HIGUCHI, K. and OKAZAKI, J. 2004. Marra Mamba ore, its mineralogical properties and evaluation for utilization. ISIJ International, vol. 45. pp. 432. POVEROMO, J.J. 2005. IMAR 7: Industrial rocks and minerals, 7th ed. Society for Mining, Metallurgy, and Exploration. 1568 pp.
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The Journal of The Southern African Institute of Mining and Metallurgy
The recovery of manganese products from ferromanganese slag using a hydrometallurgical route by S.J. Baumgartner* and D.R. Groot† Paper written on project work carried out in partial fulfilment of B. Eng. (Metallurgical Engineering)
The ferromanganese industry is under pressure to deal with the slag arising from the production of ferromanganese, which is discarded in landfills or slag heaps. This material poses an environmental and health risk to surrounding ecosystems and communities, and disposal costs are increasing. Ferromanganese slag contains an appreciable amount of residual manganese metal, which can be exploited. Previous work has shown that the slag can be leached fully, while rejecting the silica to a residue. The methods that were investigated to recover manganese from the leach solution included hydroxide precipitation to upgrade the leach solution followed by manganese carbonate precipitation to produce a pure manganese carbonate product or a manganese carbonate furnace feed material, which would be recycled to increase manganese recoveries in the production of ferromanganese. In addition, electrowinning of electrolytic manganese dioxide from the leach solution was studied. The methods were compared in terms of selectivity, costs, and product quality. Co-recovery of the leach residue, which is a potential cement additive, is discussed. Among the methods investigated to upgrade the pregnant leach solution, hydroxide precipitation utilizing ammonia to adjust the pH appears to be the most effective in removing major impurities such as iron, aluminium, and silica to less than 1 ppm. The manganese carbonate and impure manganese carbonate furnace feed products met quality specifications. However, although the production of these materials was technically viable, the large amounts of base reagent that were required to raise the pH, and the associated high operating costs, rendered the process uneconomic. An optimization study was therefore carried out with the primary objective to determine the ideal acid amount to be utilized in the waterstarved digestion stage, thereby decreasing acid and base consumption while optimizing the quality of the pregnant leach solution, and producing a leach residue that contained <1% Mn. The outcome was an economically viable process. Additional benefits included an increase in the manganese content of the impure manganese carbonate furnace feed material, and a substantial reduction in the dilution of the pregnant leach solution, thereby maintaining high manganese concentrations that rendered the solution viable for electrowinning of electrolytic manganese dioxide, the production of which yielded a current efficiency of 74%. Keywords environment, ferromanganese slag, manganese products, water-starved digestion, precipitation.
Introduction A high manganese recovery is vital in rendering the ferromanganese production process economically viable. It is particularly crucial in recovering manganese products from The Journal of The Southern African Institute of Mining and Metallurgy
* Palabora Copper Ltd, South Africa. † Department of Materials Science and Metallurgical Engineering, University of Pretoria, Pretoria, South Africa. © The Southern African Institute of Mining and Metallurgy, 2014. ISSN 2225-6253. Paper received Jan. 2014. VOLUME 114
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Synopsis
secondary sources such as ferromanganese slag. This industrial waste material contains an appreciable amount of residual manganese that can be economically exploited. Ferromanganese slag is a waste product from the production of ferromanganese metal in blast furnaces and submerged arc furnaces. There is an estimated 53 Mt of slag dumped in South Africa (Parker and Loveday, 2006). Discarding the slag not only entails disposal costs, but also poses an environmental and health threat, mainly due to the leaching of mobile manganese in the slag. The slag can be fed back into the furnace. This option, however, is not favoured due to the build-up of alkali metals and zinc in the smelting process. The slag can also be crushed and ground to produce cement aggregate or slag cement. The disadvantage to this, however, is that there is a danger that manganese can leach out, and only freshly tapped slag can be utilized. A hydrometallurgical route that can be adopted to recycle both the discarded and freshly tapped slag would be advantageous. The slag would be subjected to a water-starved digestion and leached to produce a pregnant leach solution from which manganese could be recovered using various methods, including precipitation and electrowinning, and a residue that can be sold as an additive for Portland cement. The research focus was therefore aimed at developing economically and technically viable processes to produce a pregnant leach solution with a high manganese content and low impurity concentration. The pregnant leach solution could be utilized to produce saleable products that meet the quality requirements, or products containing manganese in a low oxidation state that could be fed back into the
The recovery of manganese products from ferromanganese slag ferromanganese production process as a sweetener, thus circumventing the environmental and health impacts of processing ferromanganese slag. In addition, a saleable byproduct such as a cement additive would be produced.
Leaching of ferromanganese slag Various authors have undertaken prior investigations into the leaching of ferromanganese slag. The leaching processes discussed include ferric chloride in the presence of sucrose, carbamate leaching, and the quick leach (also known as the water starved digestion method) with concentrated sulphuric acid (98–99%). Naganoor et al. (2000) used the ferric chloride route to leach slag. It was deduced that roasting prior to leaching was essential to ensure that the manganese was converted into a soluble form. A manganese recovery of 82% was achieved in 2 hours at a temperature of 80˚C. The oxidation of sucrose, however, increased the rate of leaching, yielding an 86% recovery in 1 hour. The pregnant leach solution can be further treated to produce electrolytic manganese dioxide (EMD) and electrolytic manganese metal (EMM). Mcintosh and Baglin (1992) investigated the feasibility of using ammonium carbamate as a reagent to recover manganese from slag. A recovery of 99% was achieved in 3 hours at a temperature of 65˚C. The manganese carbamate can be added as a sweetener in the steelmaking industry or can be further purified. The slag that was used to produce the pregnant leach solution in previous investigations was first crushed and milled, then premixed with deionized water followed by water starving (digesting) the slurry in concentrated sulphuric acid (98%) (Groot et al., 2013). A solidified cake was produced, which was left to age for 24 hours; this resulted in an increase in manganese recovery. Thereafter, the cake was water-leached in a 1–3 stage water leach, recovering between 75–95% of the manganese into solution. These results, however, were dependent on the number of leaching stages.
Recovery of manganese as manganese carbonate Zhang et al. (2010) investigated carbonate precipitation of manganese from a pregnant leach solution at 60°C using sodium carbonate (Na2CO3 ) at different pH values. The experiments were conducted under ambient conditions, but with slow agitation to reduce oxidation by air. It was observed that at pH>7.5 a 90% manganese recovery was obtained with co-precipitation of calcium and magnesium at recoveries of 43% and 13%, respectively. At pH>8.5, 99.5% of the manganese was recovered, with co-precipitation of 97% of the calcium. According to Zhang et al. (2010) the precipitation of manganese carbonate is thermodynamically favoured over magnesium carbonate and (but to a lesser degree) calcium carbonate. This is attributed to the smaller log K(Ca/Mn) value measured in comparison to the theoretical K value. This is an important consideration when attempting to recover manganese carbonate that meets specifications from a solution that contains high magnesium and low calcium concentrations. The results obtained by Zhang et al. suggested that the recovery of manganese via a carbonate precipitation yielded a good quality product. This, however, was dependent on the manganese to magnesium ratio.
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An alternative process to produce manganese carbonate (MnCO3) from a manganese solution was developed by INCO for the recovery of metal values from ocean-floor nodules in the presence of 3–8% NH3 and 1–6% CO2. The reagent acts both as a lixiviant and as a precipitant for manganese as MnCO3, which is recovered in the residue; nickel, cobalt, and copper are solubilized as stable ammines (Illis and Brandt, 1975; Zhang and Cheng., 2007). Kono et al. (1986) conducted similar studies into the recovery of manganese from nodules using a sulphurous acid leach, with precipitation of manganese as MnCO3 in the presence of (NH4)2CO3 while copper, nickel, and cobalt remained in the solution as ammine complexes.
Production of electrolytic manganese dioxide (EMD) Electrolytic manganese dioxide (EMD) is a vital ingredient in Leclanche-type dry cells. Owing to modern advances in the electronics industry, which have necessitated a greater capacity for manganese dioxide production, natural manganese dioxide (NMD) has been replaced by electrolytic manganese dioxide (EMD) (Rethinaraj and Visvanathan, 1993). The half-cell reactions are given below: [1] [2] [3] The current efficiency at the anode is affected by the following parasitic reactions (Te Riele, 1983): The evolution of oxygen: [4] The oxidation of Mn2+ to Mn3+, which diffuses into the electrolyte solution: [5] Mn3+ is an unstable ion that disproportionates to produce Mn2+ and MnO2, which forms sludge: [6] An additional parasitic reaction that can occur if there is an appreciable amount of iron in solution, and which is responsible for decreasing current efficiencies and contaminating the final EMD product, is given by: [7] Table I shows the effect that different concentrations of iron in solution have on the final product (Te Riele, 1983).
Table I
Effect of iron content of solution on iron contamination in the EMD (Te Riele, 1983) Fe in electrolyte (ppm) 0–10 10–50 45–80 60–90
Fe in final EMD (%) 0.021 0.098 0.149 0.19
The Journal of The Southern African Institute of Mining and Metallurgy
The recovery of manganese products from ferromanganese slag There are several factors that affect the quality of EMD produced, one of which is the acid concentration of the electrolyte. A high acid concentration (>50 g H2SO4 per litre) has a detrimental effect on the amount of battery-active EMD plated. In addition, the morphology is affected, resulting in a less compact deposit, and current efficiency is decreased by the formation of Mn3+ which is promoted by the higher acid concentration (Te Riele, 1983). Temperature is another critical factor. A temperature above 90°C results in a desirable current efficiency, which ranges between 90–97%, provided the system is operated at 100 A/m2 (Te Riele, 1983; Rethinaraj and Visvanathan, 1992). At lower temperatures, however, the current efficiency will be variable depending on the type of materials utilized for the electrodes, and the cell voltage will increase, thereby increasing energy consumption. In addition, lower temperatures result in shorter discharge times on the EMD battery cell (Te Riele, 1983).
Experimental procedure The flow diagram in Figure 1 summarizes the experimental routes that were investigated to determine the most technically and economically viable methods of producing various manganese products.
Materials and methods Materials A South African ferromanganese producer in KwaZulu-Natal supplied the ferromanganese slag utilized in previous investigations. The chemical and phase compositions of the slag were determined using X-ray fluorescence spectrometry (XRF) and X-ray diffraction spectrometry (XRD), which were performed by Scrooby's Laboratory Service and the Department of Geology at the University of Pretoria.
Chemical composition of the ferromanganese slag The typical composition of the slag, determined by XRF analysis, is given in Table II.
The major elements were manganese, silicon, and calcium. The iron content was low, which was ideal as iron results in difficulties in the downstream purification stages. The major phases present in the material included glaucochroite, manganosite, and gehlenite as shown in Table III. The manganese content of the slag (approx. 30%) was sufficient for the investigation to proceed further. All reagents were of analytical grade, and supplied by Merck Millipore, and distilled water was used throughout. All glass and metal ware was decontaminated with detergent and rinsed thoroughly with distilled water.
Sample preparation The work was performed on 100 g ferromanganese slag samples that had been pre-crushed to 45 mm. The material was further reduced in size using a gyratory crusher, then milled to a cut size of 600 μm utilizing ball milling, and dryscreened.
Purification of the manganese pregnant leach solution The objective of the purification stage was to obtain a pregnant leach solution (PLS) that contained high manganese and low impurities concentrations, from which manganese products that meet the quality specifications could be produced.
Table II
Typical composition of the ferromanganese slag Element
Mass %
C Mn P Si Cr Ni Cu Al V Ti Co Ca Mg Fe
1 28 0.01 21 <0.005 <0.005 0.006 <0.005 <0.005 0.14 <0.005 19 3 1
Table III
Phase
Figure 1—Flow diagram illustrating the basic process routes investigated for the recovery of manganese products The Journal of The Southern African Institute of Mining and Metallurgy
Glaucochroite Manganosite Gehlenite Monticellite Quartz Amorphous
Chemical formula
Mass %
CaMnSiO4 MnO Ca2Al[AlSiO7] CaMgSiO4 SiO2 -
55.00 – 65.10 2.28 – 4.50 4.00 - 8.49 0.00 – 2.00 0.00 – 5.10 21.00 – 30.0
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Phase analysis of the ferromanganese slag (Groot et al., 2013)
The recovery of manganese products from ferromanganese slag The reagents used to increase the pH for impurity removal included sodium hydroxide (NaOH) or ammonia (NH3), which were employed in separate purification investigations. The oxidant was hydrogen peroxide (34% H2O2), to oxidize Fe2+ to Fe3+ so that iron was precipitated within a pH window of 3–5.5. A PLS sample size of 100 ml was used throughout experimental runs. Before base reagent was added, 34% H2O2 was added dropwise until the solution turned from pink to a mustard colour. For purification using NaOH, 24 g of NaOH pellets were weighed, and dissolved in 200 ml distilled water to obtain a solution of 120 g/l of NaOH, which was added dropwise to the solution to increase the pH from less than 0 to approximately 5.6. The solution was kept at room temperature and continuously stirred. A similar procedure was followed for the experiments that involved the upgrading of the PLS using 25% ammonia solution, which was added dropwise to increase the pH to approximately 5.6. The upgraded PLS was analysed for major elements by inductively coupled plasma – optical emission spectrometry (ICP-OES).
Production of manganese products Precipitation of manganese carbonate A 100 ml sample of PLS was used for the investigations utilizing Na2CO3 or (NH4)2CO3 as base reagents. The Na2CO3, at 150 g/l, was added dropwise to the PLS; the (NH4)2CO3 was added to the PLS in its solid form. Approximately 15 g of (NH4)2CO3 was added to adjust the pH from approximately 5.5 to 8.5, the pH range in which >99% of the manganese was recovered. The system was open to the atmosphere, and the procedure was conducted at room temperature (approx. 25°C) while the solution was agitated. Once precipitation was complete, the product was filtered, washed with distilled water, and dried at 30°C. The manganese and impurities contents of the barren solution obtained after precipitation were determined by ICPOES. The precipitate was analysed using XRF and XRD.
Results and discussion Pregnant leach solution and residue To produce the pregnant leach solution and residue the milled slag material was subjected to a water starved digestion and thereafter water-leached for two hours. Table IV shows a mass balance giving the general recoveries of manganese and impurity elements to the PLS and residue respectively. The mass balance was based on a 200 g slag sample. The discrepancies in the percentage recovery of elements were probably due to the inhomogeneity of the ferromanganese slag sample. It was assumed that the chemical composition of the slag was as given in Table IV.
Purification of the pregnant leach solution (PLS) Hydroxide purification It was essential to obtain a purified PLS in order to produce MnCO3 and EMD that adhered to the chemical specifications. The PLS had to contain >25 g/l Mn for the production of MnCO3, and 27–66 g/l Mn for the production of EMD (Te Riele, 1983). In addition, the PLS should adhere to the specifications in Table V. During the addition of the base reagent, oxygen was bubbled through the solution at 60°C to oxidize iron from the Fe2+ to the Fe3+ state, thus allowing iron to precipitate as a
Production of a furnace feed material (impure MnCO3) The apparatus and reagents used to produce the furnace feed material were the same as for the purification of the PLS and the pure MnCO3 product, except no filtration step was included. The product was dried in an oven at 30°C.
Figure 2—Experimental set-up for the production of EMD
Electrowinning of electrolytic manganese dioxide (EMD)
Mass balance and recovery of elements to the leach residue and PLS
The electrodes were graphite rods with a diameter of 7.5 mm. A current density of 100 A/m2 was adopted, with a plating time of 72 hours to allow a deposit thickness of approximately 2.4 mm, which would produce a mass of EMD >3 g, and a temperature of 90°C. The electrolyte volume was 1.1 litres. The set-up of the electrowinning experiment for the production of EMD is shown in Figure 2. The EMD was analysed for gamma and alpha manganese dioxide (MnO2) by XRD, and an elemental analysis was carried out by XRF. The manganese content of the solution after electrowinning was analysed using ICP-OES.
Element Slag (feed), g Residue (g) Leach solution (g) % Recovery
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Table IV
Mn P Si Cr Ni Cu Al V Ti Co Ca Mg Fe
55.4 0.02 40.8 0.012 0.274 38.2 6.34 1.86
9.344 13.120 0.346 0.057 21.71 1.872 0.330
31.7 0.115 0.034 0.025 2.927 0.367 6.307 1.205
74.1 >100 32.2 >100 >100 >100 20.8 57.8 >100 82.5
The Journal of The Southern African Institute of Mining and Metallurgy
The recovery of manganese products from ferromanganese slag hydroxide below a pH of approximately 5.6. However, this was ineffective due to the slow kinetics of dissolution of oxygen in the leach solution, and was further hindered by the solution pH < 3 (Zhang and Cheng, 2007). The oxygen was replaced with hydrogen peroxide, which was effective in oxidizing iron to the ferric state; therefore, all test work was carried out using hydrogen peroxide as the oxidant at room temperature where applicable. A window of optimal precipitation of the major impurities, including Fe, Al, and Si, lay in a pH range of 2.5 and 5.6–6. Complete precipitation of impurities occurred between pH 5.5–6. From the results in Table VI, both NaOH and NH3 were effective reagents in removing impurities from solution. However, NH3 appeared to be the more effective, resulting in a purified PLS with Fe, Al, and Si contents of less than 1 ppm, in contrast to NaOH where the solution
Table V
Standard specifications for a viable manganese PLS Typical PLS specifications, ppm Ni Co Fe Si Cu Zn
1 0.3 15 10 5 10
contained 604 ppm Fe, 8 ppm Si, and 140 ppm Al after precipitation.
Manganese products Precipitation of manganese carbonate Manganese precipitated from the purified PLS at pH >6, and manganese recovery reached 99.8% and 98.9% using (NH4)2CO3 and Na2CO3, respectively, at a pH >8.5. The results of the ICP-OES analyses before and after carbonate precipitation are given in Table VII. The results agreed with those obtained by Zhang et al. (2010), who achieved a manganese recovery of >99.5% pH >8.5. The results in Table VIII indicate that using either (NH4)2CO3 or Na2CO3 as the precipitating reagent resulted in a MnCO3 product that meets the quality specifications, with the exception of the calcium content, which exceeded the maximum allowable amount. This finding is in agreement with Zhang et al. (2010), who argued that carbonate precipitation is more selective for magnesium than calcium, due to the thermodynamics favouring the co-precipitation of calcium, which is attributed to the low log K(Ca/Mn) value determined in comparison to the theoretical K value. The lower MnCO3 yields from Tests 1, 3, and 5 were due to the co-precipitation of additional phases such as Na2CO3 and (NH4)2Mg(SO4)2.6H2O. These phases, however, are water-soluble; therefore the product was washed with water to yield products that contained >92% MnCO3.
Table VI
Element concentrations in the PLS before and after hydroxide precipitation with NH3 or NaOH Species
Units
Reagent NaOH
Mn Cr Fe K Mg Al Al Na Na Ni Si Zn
g/L ppm ppm ppm g/L ppm g/L ppm g/L ppm ppm ppm
NH3
PLS pH ~-0.3
PLS pH ~5.6
PLS pH ~-0.3
PLS pH ~5.6
36 40 1568 117 7 3 366 21 157 11.3
19 <1 17.8 4 3 56 56 7 <1 <1
24 20 842 240 4 2038 115 11 <1 9
15 <1 <1 262 2 <1 86 5 <1 <1
Table VII
Chemical composition of the PLS before and after carbonate precipitation Unit
Reagent Na2CO3
Mn Mn Ca Fe Mg
g/L ppm ppm ppm ppm
(NH4)2CO3
PLS pH ~5.6
PLS pH ~8.5
PLS pH ~5.6
PLS pH ~8.5
18.65 209.1 17.8 3437.50
650.3 10.8 9.4 643
14.76 181.6 <1 2321.69
4.8 14.1 <1 1608.70
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Species
The recovery of manganese products from ferromanganese slag Table VIII
Chemical composition of the MnCO3 produced and the standard specifications (www.alibaba, 2013) Parameter
Standard specification (%)
Test 1 (%)
Test 2 (%)
Test 3 (%)
Test 4 (%)
Test 5 (%)
>92 >43 0.08% max. 0.1% max. 0.1% max. 0.1% max. 1% max.
56.48 44 0.31 1.02 -
98.7 43.5 <0.01 1.27 -
94.15 42.6 1.02 2.34 -
98.1 43 0.93 1.86 -
79.54 44 0.02 1.91 -
MnCO3 Mn Iron as Fe Nitric acid-insoluble Chloride as Cl Ca Sulphate as SO4
A disadvantage of using Na2CO3 as a base reagent in this system was that between pH 7.3–7.5 buffering appeared to have occurred, due to the buffering action of the bicarbonate that formed from the dissociation of carbonic acid. To overcome this, a stoichiometric amount of NaOH was added to increase the pH to approximately 8.5. After the addition of the NaOH the pH was raised to >8.5, resulting in 98.9% Mn recovery, in comparison to 96% Mn recovery at pH 7.5.
Furnace feed material The production of the MnCO3 furnace feed material involved the precipitation of impurities followed by the precipitation of manganese as MnCO3, without a filtration step. As in the production of pure MnCO3, the manganese recovery reached >99% using (NH4)2CO3 or Na2CO3 as the base reagent at pH >8.5. The XRD and XRF results are given in Tables IX and X. These results suggest that using NaOH and Na2CO3 or NH3 and (NH4)2CO3 resulted in a product that can potentially be fed into a furnace (>35% Mn), as shown in Tests 1 and 3 in Table X. The MnCO3 furnace feed produced by the NaOH and Na2CO3 precipitation method contained appreciable amounts of sodium (approx. 13 %), which would be detrimental to furnace operations because of the build-up of alkali metals in the furnace, which may lead to explosions and refractory attack. With the use of NH3 and (NH4)2CO3, approximately 45% of the product contained ammonium phases, which would pose a problem to furnace operations owing to the dissociation of the ammonium phases to NH3 and the further dissociation of ammonia to H2 and N2 at temperatures above 400°C. The ammonium and sodium phases are water-soluble; to rid the material of these phases a water wash step was employed in order to render the product viable as a furnace feed.
Optimization study An optimization study was undertaken with the primary objective of decreasing the base reagent addition required to neutralize the PLS by investigating the ideal acid amount to be used. The benefits of this included a reduced acid and base reagent consumption, which resulted in a reduction in
Table IX
XRD results for the MnCO3 furnace feed material produced via NaOH and Na2CO3, or NH3 and (NH4)2CO3 precipitation Test
Reagent
1
NaOH and Na2CO3
2
NaOH and Na2CO3
3
NH3 and (NH4)2CO3
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MnCO3
n.d.
Na2SO4
n.d.
(Ca,Mn)(Ca,Mn,Fe,Mg)(CO3)2
n.d.
MnCO3
60.26
Na2SO4
39.74
MnCO3
56.48
(NH4)2Mg(H2O)6(SO4)2
35.88
NH4SO4
7.64
Table X
Standard specifications (Gous, 2013) and results of tests carried out to produce MnCO3 furnace feed Precipitation method
Specification (%)
MnCO3
To determine the net present values (NPVs), it was decided to utilize past and present production and plant history of the supplier of the slag. In addition, the capex and opex values were calculated based on 2013 expenditures. The NPVs for the products investigated were negative, due to the high acid content of the PLS, which required a large base reagent input to neutralize the solution which in turn increased the operating costs.
Mass %
n.d. A quantitative analysis could not be carried out
Parameter
Economic analysis
Phase
NaOH and
NH3 and
Na2CO3
(NH4)2CO3
Test 1 (%)
Test 2 (%)
Test 3 (%)
-
-
60.26
56.48
Mn
>35
44
27.7
36.2
Fe
-
0.31
1.22
0.9
Na
<0.8
#
12.9
*
K
<0.9
*
*
*
Zn
<0.07
Trace amounts
Trace amounts
Trace amounts
# Not analysed for
* Element not contained in product
The Journal of The Southern African Institute of Mining and Metallurgy
The recovery of manganese products from ferromanganese slag operating costs, thereby resulting in a positive net profit value and rendering the production processes economically viable. In addition, the technical viability was improved as the amounts of impurities in the leach solution were reduced, as most were precipitated and reported to the residue. This was beneficial for the production of the impure MnCO3 furnace feed material, which had a higher manganese content and lower impurities, thereby yielding a product with a high Mn:Fe mass ratio, which was ideal. In addition, due to the reduction in base reagent consumption, dilution of the PLS was reduced and a high manganese concentration was maintained. This was beneficial for the electrowinning of EMD, since high manganese concentrations (>27 g/l) were required. The results of the study are summarized in Figures 3 and 4. From the results the ideal acid addition to produce a PLS with high manganese and low impurities content appears to be 40 ml; the pH at this amount is approximately 4.3 and an appreciable amount of impurities had precipitated out, as is evident in Figure 4. The concentrations of Al, Fe, and Si obtained were 5 ppm, 684 ppm, and 39 ppm respectively, in contrast to an acid addition of 100 ml, which resulted in Al, Fe, and Si concentrations of 4 g/l, 1.6 g/l, and 157 ppm. This reduced the Al, Fe, and Si contents of the PLS by 99.88%, 42.8% and 75%, respectively, compared with the original water-starved digestion acid amount of 100 ml. The manganese concentrations at 40 ml acid addition were > 29 g/l, which rendered the leach solution amenable to the production of quality manganese products and was suited for the electrowinning of EMD.
The manganese concentration appeared to decrease with acid additions of less than 50 ml. This was attributed to incomplete reaction of the slag, as less than 50 ml is below the stoichiometric amount, and thus the silicate matrices could not be completely destroyed. The slight decrease in manganese concentration between 40–60 ml acid was probably due to material variance; this is further substantiated by the results given in Table XI, which clearly illustrate the difference between the tests conducted with 50 ml acid under identical conditions and with the same slag sample. The manganese concentrations obtained in test 1 and test 2 differed by 6 g/l, and variances existed between the major impurity elements also. A countercurrent three-stage water wash was employed on the material produced after the slag material was digested using a 40 ml acid addition. It was vital that the manganese content of the residue was brought below 1% for it to be a viable cement additive. Figure 7 illustrates the extent to which manganese was leached from the material using 40 ml of acid. From the results in Figure 5 it is evident that the threestage water wash was able to reduce the manganese content of the residue to approximately 2.7 %, which did not meet the specification of < 1% manganese, thus a 40 ml acid addition was not ideal. A further study was therefore carried out to determine the optimum acid addition required to render the residue suitable for a quality cement additive containing < 1% Mn. From the results in Figure 6 it appears that an acid addition of 50—55 ml would be ideal, and would yield a manganese content of less than 1% in the residue while still producing a pregnant leach solution of pH 3.5–4 (Figure 3), from which products with a high manganese content can be produced. Table XI
Example of elemental variance in the PLS at 50 ml acid volume in the water-starved digestion Acid volume 50ml Element
Figure 4—Effect of acid addition on the concentrations of Mn, Mg, Al, Fe, and Si in the PLS The Journal of The Southern African Institute of Mining and Metallurgy
Test 2
30 g/L 313 ppm 917 ppm 6 g/L
36 g/L 359 ppm 883 ppm 7 g/L
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Figure 3—pH of the pregnant leach solution as a function of acid addition
Mn Al Fe Mg
Test 1
The recovery of manganese products from ferromanganese slag Table XII
Chemical composition of the PLS using 40 ml acid addition in the water-starved digestion Major element analysis on PLS – 40 ml acid volume
Figure 6—Manganese content in the leach cake as a function of acid amount employed in the water-starved digestion
Production of a furnace feed material The pH of the PLS utilized to produce the impure manganese carbonate furnace feed product was approximately 4.3. The manganese and impurities contents are given in Table XII. Since reducing the acid amount from 100 ml to 40 ml resulted in an increase in pH to 4.3, it was decided to add Na2CO3 or (NH4)2CO3 to increase the pH to >8 in order to precipitate manganese as MnCO3. However, phases such as (NH4)2Mg(SO4)2.6H2O and Na2SO4 co-precipitated and contaminated the product, rendering it unsuitable as a furnace feed. The product was therefore subjected to a washing step with distilled water, as these phases are watersoluble, in an attempt to produce a viable furnace feed. A pure manganese carbonate (>92%) resulted; the XRF analysis is given in Table XIII. The resulting manganese carbonate precipitates yielded >95% MnCO3, with >45% Mn content, and contained less than 9% impurities, of which iron constituted 1–1.2%. Manganese and iron co-precipitated at a pH of approximately 7 when iron was not oxidized and carbonate was added to the PLS (Pakarinen, 2011). As shown by Figure 7, the formation of FeCO3 is impossible due to the high redox potential; however, iron will most likely precipitate as ferric oxohydroxide (FeO(OH)). It is evident from the results that the composition of the furnace feed material was improved by increasing the pH. Furthermore, the water-soluble phases were effectively removed by washing with water. It can be concluded that the process conditions resulted in a technically viable product. The XRF analyses in Table XIII indicate that a product with high manganese content and high Mn:Fe ratio of approximately 48:1 was obtained, which acceptable as an additive in ferromanganese production.
Element
Unit
Concentration
Mn Al B Ca Cr Fe K Mg Na Ni Si Sr V Zn
g/l ppm ppm ppm ppm ppm ppm ppm ppm ppm ppm ppm ppm ppm
29.85 6 240 637 <1 663 104 5411 143 7 52 9 <1 2
Figure 7—Eh-pH diagram for manganese and iron compounds (Pakarinen, 2011)
Table XIII
XRF analyses of impure MnCO3 furnace feed produced using (NH4)2CO3 or Na2CO3 as base reagents Base reagent
Economic outcome
Element (%)
To render the production of the various manganese products viable, changes were made to the water starved digestion process after an optimization study had been undertaken to determine the ideal acid amount that would be required to decrease the base consumption. It was determined that an acid amount of 50 ml would be ideal to achieve this. The economic outcome of this alteration is that it rendered the NPVs of the various manganese products positive. In addition, the change to acid amount yielded an impure MnCO3 furnace feed material with a manganese content of > 45%, which was higher than the ideal manganese content
Mn P Si Cr Ni Cu Al V Mo Co Ca Mg Fe
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(NH4)2CO3
Na2CO3
48.8 <0.005 <0.005 0.09 <0.005 <0.005 <0.01 <0.005 <0.005 <0.005 4.43 1.71 1.13
47.3 <0.005 <0.005 0.09 <0.005 <0.01 <0.01 <0.005 <0.005 <0.005 4.83 3.72 1.07
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The recovery of manganese products from ferromanganese slag of between 35–40%, which is sought after by the ferroalloy industry. The benefits of the feed material produced via the hydrometallurgical route include a high quality feed material which will result in an increase in the production of ferromanganese metal per day as the recoveries of manganese will be increased, and thereby industry will have the option of purchasing lower grade ore (high Fe ore) at a reduced cost (Steenkamp, 2013).
Electrowinning of EMD We decided to electrowin electrolytic manganese dioxide (EMD) from the PLS by water-starved digestion of the slag with 40 ml of sulphuric acid. The cake was then waterleached with 350 ml of distilled water to ensure that the manganese concentration in the leach solution was >27 g/l, allowing it to be amenable to electrowinning of the EMD product. This method was employed due to the chemical variance of the slag sample, and to ensure that the recovery of manganese to the leach solution was >27 g/l. The concentrations of the manganese and impurity elements in the leach solution before and after purification with ammonia solution are given in Table XIV. The mass of EMD produced over a 72 hour period was 8.3 g, thus yielding a current efficiency of approximately 74%. A current density of 100 A/m2 should result in current efficiencies that typically lie between 90–97%, as was achieved by by Te Riele (1983). The current efficiency in this investigation, however, may have been affected by the increase in acid concentration, which promoted the formation of Mn3+. This can be overcome by increasing the pH by
adding a base reagent such as ammonia, in order to increase the pH to 5–6, thereby maintaining an ideal current efficiency. The results in Table XV indicate that a product that meets specifications can be produced. The iron content was high; however, this result is not in agreement with the ICP-OES analysis of < 1 ppm Fe (Table XIV). According to Te Riele (1983) a pregnant leach solution containing < 1 ppm Fe should yield a product with only 0.021 % Fe. Furthermore, SEM-EDS analysis of the material did not detect iron. The micrograph in Figure 8 clearly show slight pitting of the surface of the EMD. Figure 9 is considered to show the typical morphology of EMD. The morphology appears to be similar to that obtained by Liu et al. (2007). Thus, good quality EMD can be produced from the leaching of ferromanganese slag.
Conclusions and recommendations The manganese content of the slag (>30%) rendered the material attractive for further processing. A pregnant leach solution with typical manganese concentration of 25–35 g/l and low impurities content was obtained. Impurities were effectively removed using ammonia or sodium hydroxide. Ammonia was the most effective reagent, removing Fe, Si, and Al to levels of <1 ppm.
Table XIV
Chemical composition of the PLS used for electrowinning of EMD before and after purification with NH3 Species
Unit Before purification pH ~ 4.5 After purification pH ~5.6
Mn Cr Fe K Mg Al Na Ni Si Zn
g/L ppm ppm ppm g/L ppm ppm ppm ppm ppm
30.8 1 674 162 5.6 81 181 6 79 2
Figure 8—SEM micrographs of the surface of the EMD deposit at 500 μm and 250 μm resolution
30.8 <1 <1 234 5.1 <1 237 6 22 <1
Table XV
Analysis of the EMD product
MnO2 Mn Fe Al Cu Pb
Battery-active EMD specifications
Results
91%
> 95%
>59% <10 ppm <100 ppm <5 ppm < 5 ppm
68% 700 ppm 5 ppm < 1 ppm Trace
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The recovery of manganese products from ferromanganese slag Water-starved digestion of the slag with 100 ml sulphuric acid and purification of the resulting solution led to a manganese recovery of >99.7% into a manganese carbonate product at a pH >8.5. However, co-precipitation of calcium carbonate resulted in the product exceeding the maximum specified calcium content. Further investigations will be aimed at investigating the kinetics of co-precipitation in order to decrease the amount of calcium carbonate precipitated. A viable MnCO3 furnace feed material with a high manganese (> 45% Mn) and low iron content (< 2% Fe) was produced from the pregnant leach solution utilizing NH3 and (NH4)2CO3 or NaOH and Na2CO3 to increase the pH to >8.5. Water-soluble phases containing ammonia and sodium that formed could be detrimental to furnace operations. To rid the feed material of these phases and to yield a quality product the material was washed with water, which yielded a product containing >95% MnCO3. Although the processes were shown to be technically feasible, the large amounts of the base reagent required to adjust the pH of the pregnant leach solution from less than 0 to 5.6 when 100 ml of sulphuric acid was utilized in the digestion stage resulted in an increase in operating costs, rendering the production of the manganese products unviable. Therefore, to improve economic viability a study was undertaken to optimize the quality of the PLS, and to produce a leach residue (cement additive) that contained < 1% Mn. It was found that decreasing the sulphuric acid addition to approximately 50 ml resulted in an increase of the pH of the PLS to approximately 3.4–4, and within this pH range most of the major impurity elements precipitated out and reported to the residue. This adjustment resulted in the production of a MnCO3 furnace feed material with a manganese content >45% and a high Mn: Fe ratio. In addition, to produce pure manganese carbonate, a smaller amount of neutralizing reagent would be required to adjust the pH to 5.6 when utilizing an acid addition of 50 ml. This adjustment rendered the process economically viable. Additional benefits of using 50 ml of acid in the digestion stage include a significant reduction in the dilution of the pregnant leach solution with large amounts of base reagent, thereby maintaining a high manganese concentration in the leach solution, and eliminating potential gelling problems. The use of a hydrometallurgical process to recover manganese products and produce a leach residue that is saleable as a cement additive is beneficial to the environment and surrounding communities in ferromanganese-producing areas. Both slag dumps and current arisings can be utilized, thereby reducing the environmental and health risks posed by the leaching of mobile manganese contained in the slag. In addition, all potential waste streams produced during the hydrometallurgical process, such as the barren solution, can potentially be recycled to produce additional marketable byproducts, including ammonium sulphate or sodium sulphate. However, further investigations would have to be carried out to confirm this.
GROOT, D., KAZADI, D., POLLMANN, H., DE VILLIERS, J., REDTMANN, T., and STEENKAMP, J. 2013. Utilization of ferromanganese slags for manganese extraction and as a cement additive. Advances in Cement and Concrete Technology in Africa, Emperor's Palace, Johannesburg, 28– 30 January 2013. pp. 984–985. GOUS, J. 2013. Transalloys. Personal Communication. ILLIS, A. and BRANDT, B.J. 1975. Selective process for the recovery of metal values from sea nodules. CA Patent no. 974371. KONO, Y., MIZOTA, T., and FUJII, Y. 1986. A precipitation separation method for copper, nickel, and cobalt recovery from sulfurous acid leach liquor of sea manganese nodules. Nippon Kogyo Kaishi, vol. 102, no. 1183. pp. 585–590. LIU, B., THOMAS, P.S., RAY, A.S., DONNE, S.W., and WILLIAMS, R.P. 2007. DSC characterisation of chemically induced electrolytic manganese dioxide. Journal of Thermal Analysis and Calorimetry, vol. 88. pp. 177–180. MCINTOSH, S.N. and BAGLIN, E.C. 1992. Recovery of manganese from steel plant slag by carbamate leaching. Report of investigations, US Department of the Interior, Bureau of Mines. NAGANOOR, P.C., PRASANN, A.S.R., SHIVAPRASAD, K.H., and BHAT, K.L. 2000. Extraction of manganese from ferro-manganese slag. International Symposium on Processing of Fines, Jamshedpur, India, 2–3 November 2000. National Metallurgical Laboratory, Jamshedpur. pp. 300–306. PAKARINEN, J. 2011. Recovery and refining of manganese as by-product from hydrometallurgical processes. Colloqium, Lappeenranta University of Technology, Finland. p. 29. PARKER, J. and LOVEDAY, G. 2006. Recovery of metal from slag in the ferro-alloy industry. Hidden Wealth. Southern African Institute of Mining and Metallugy, Johannesburg. pp. 7-15. RETHINARAJ, J.P. and VISVANATHAN, S. 1993. Preparation and properties of electrolytic manganese dioxide. Journal of Power Sources, vol. 42. pp 335–343. STEENKAMP, J. 2013. Department of Materials Science and Metallurgical Engineering, University of Pretoria. Personal Communication. STEENKAMP, R. 2013. Exxaro. Personal Communication TE RIELE, W.A.M. 1983. Electrowinning of manganese dioxide. SAIMM Vacation School Electrometallurgy, Randburg. South African Institute of Mining and Metallugy, Johannesburg. pp. 238-261. WELLBELOVED, D.B, CRAVEN, P.M., and WAUDBY, J.W. 2003. Ullmann's Encyclopedia of Industrial Chemistry. Wiley and Sons, New York. Chapter 6. ZHANG, W. and CHENG, C. 2007. Manganese metallurgy review Part II: Manganese separation and recovery from solution. Hydrometallurgy,
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ALIBABA GROUP. 2013. Manganese product specifications and price. http://sjzbcchem.en.alibaba.com/product/517509869-
ZHANG, W., CHENG, C., and PRANOLO, Y. 2010. Investigation of methods for
213404979/sell_manganese_carbonate_MnCO3.html [Accessed 15 August
removal and recovery of manganese in hydrometallurgical processes.
2013].
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Wear of magnesia-chrome refractory bricks as a function of matte temperature by M. Lange*, A.M. Garbers-Craig*, and R. Cromarty* Paper written on project work carried out in partial fulfilment of B. Eng. (Metallurgy)
The postulation that primary platinum group metal (PGM) matte will chemically react with magnesia-chrome bricks when temperatures exceed 1500°C was tested. Magnesia-chrome brick samples were heated in contact with matte at 1300°C to 1750°C for 30 minutes, after which the refractory samples were analysed using reflected light microscopy and scanning electron microscopy. The samples were all completely penetrated by matte. As the temperature increased the matte also penetrated the fused aggregate grains and disintegrated them. The chromium concentration of the matte inside the refractory samples was found to be slightly higher than that of the bulk matte. At temperatures of 1500˚C and higher, MgO, FeO, and magnesium-rich silicate crystals could be identified in the matte directly adjacent to the refractorymatte interface. Phase relations clearly indicated that chemical reactions take place between primary PGM matte and the magnesiachrome refractory material at temperatures above 1500°C, but that these reactions are more complex than expected from FactSAGE® calculations. Keywords refractory, magnesia-chrome bricks, penetration, fused grains.
Introduction The Bushveld Complex in South Africa is the largest known layered igneous complex and contains two-thirds of the world’s PGM reserves (Jacobs, 2006). Platinum-group metals (PGMs) are concentrated in three narrow layers in the Bushveld Complex, namely the Merensky Reef, the Platreef, and the UG2 chromitite layer (Jones, 2005). PGMs occur together with base metal sulphides in the Merensky and Platreef, while the UG2 layer has a high chromite content and low concentrations of base metal sulphides. UG2 ores contain approximately 30% Cr2O3 compared with 0.1 % Cr2O3 in the Merensky deposit (Jones, 2005). Mines are increasingly exploiting the UG2 as the Merensky Reef becomes depleted. The chromite content of the UG2 ore presents a major challenge for the PGM smelting process as high levels of chromium increase the slag liquidus temperature, which necessitates higher operating temperatures. If, The Journal of The Southern African Institute of Mining and Metallurgy
* Centre for Pyrometallurgy, Department of Materials Science and Metallurgical Engineering, University of Pretoria, Pretoria, South Africa. © The Southern African Institute of Mining and Metallurgy, 2014. ISSN 2225-6253. Paper received Jan. 2014. VOLUME 114
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Synopsis
under non-reducing conditions, the Cr2O3 content of the concentrate exceeds approximately 1.8% the chromite spinels do not dissolve in the slag at conventional slag temperatures (1500–1550°C), but build up in the furnace hearth and form a ‘mushy’ threephase layer at the slag-matte interface (Eksteen 2011; Hundermark, 2011). This leads to inefficient matte segregation and difficulties in matte tapping. To achieve adequate suspension of the spinels, the furnace temperature is increased by increasing the electrode immersion depth and power density (Eksteen, 2011). Chromium control is therefore an important factor when considering furnace operating temperature. This is done by effectively controlling the concentrate blend that is fed into the furnace and adjusting the operating power accordingly (Jacobs, 2006). Of great concern when smelting concentrates of high chromite content is high matte temperatures that can destroy the protective slag freeze lining when the matte temperature is higher than the slag liquidus temperature (Eksteen 2011; Hundermark 2011; Warner et al., 2007). Slag temperatures of between 1600°C and 1700°C and matte temperatures of between 1450°C and 1500°C were estimated by Snyders et al. (2006) at the Polokwane smelter. These high operating temperatures result in highly fluid matte, which causes high tap rates and severe matte penetration into the tapping channel and endwall. Failure of the taphole bricks can cause major damage to the furnace and poses a great safety risk to operators. FactSAGE® calculations performed by Eksteen (2011) predicted that at temperatures above 1500°C the matte will react chemically with the refractory lining. In this process chromium will be picked up in the matte and
Wear of magnesia-chrome refractory bricks as a function of matte temperature oxygen will be transferred from the refractory to the matte according to reactions [1] and [2]: [1]
With ΔG ° = -8.478 kJ/mol at T=1500°C (ΔG ° < 0, when T > 1450°C) [2]
With ΔG ° = +11.03 kJ/mol at T=1500°C (ΔG ° < 0, when T > 1650°C) Eksteen (2011) concluded that the combination of high temperatures and sulphidation reactions would cause disintegration of the brick due to the destruction of its spinel bonding phase. This project subsequently investigated whether chemical interaction between magnesia-chrome bricks and primary PGM matte takes place at temperatures above 1500°C, as predicted by Eksteen (2011).
➤ In the second design, matte was contained in a closed alumina crucible. A cylindrical sample of refractory brick was inserted into the matte (Figure 1). There was a limited volume of refractory in contact with the matte, and the refractory sample and matte were isolated from the graphite susceptor. The induction furnace was purged with technical-grade argon (99.9% Ar) for 30 minutes, after which it was heated to temperatures ranging from 1300°C to 1750ºC at 25°C/min. The samples were kept at temperature for 30 minutes. The furnace was cooled down at 25°C/min under argon. Experimental runs were randomized to improve statistical accuracy. The cooled samples were visually inspected, measured, cut, and polished sections prepared. Samples were coldmounted in epoxy resin, ground, and finally polished to a 1 μm finish. The polished sections were examined using reflected light microscopy and scanning electron microscopy. SEM-EDS was used to evaluate the amount of matte penetration and the possible chemical interaction between the matte and refractory.
Results and discussion Microstructure of unreacted magnesia-chrome refractory brick
Materials Vereeniging Refractories supplied samples of reconstituted fused-grain magnesia-chrome refractory bricks, while furnace matte and slag samples were obtained from the Anglo Platinum Waterval Smelter. According to the supplier’s specifications, the brick had a chemical composition of 57.4% MgO, 21.3% Cr2O3, 11.2% Fe2O3, 7.8% Al2O3, 1.5% SiO2, and 0.8% CaO, and an apparent porosity of 16%. The matte was analysed using powder X-ray diffraction (XRD), scanning electron microscopy–energy dispersive X-ray spectroscopy (SEM-EDS), inductively coupled plasma optical emission spectrometry (ICP-OES), and sulphur analysis (Leco-S). The matte sample consisted mainly of nickel, copper, and iron sulphides, and contained 39% Fe, 15% Ni, 11% Cu, 34%S, 0.61% Cr, and 1% SiO2.
The as-supplied magnesia-chrome brick consisted of large magnesia grains (dark grey phase), which contained small amounts of iron and chrome in solid solution (Figure 2). Finely exsolved spinel crystals (light grey phase) were observed in the magnesia grains. The spinels on the grain boundaries had roughly the same composition as the spinels within the magnesia grains.
Design 1 – magnesia-chrome crucible containing matte and slag
Experimental
Design 1 (magnesia-chrome crucible, filled with matte and slag) was used for tests at 1300°C and 1500°C. This design, however, resulted in the refractory soaking up all the matte, leaving no matte in the crucible for analysis. The microstructure of the magnesia-chrome brick with soaked-up matte is shown in Figures 3 and 4.
Laboratory-scale experiments were performed in an induction furnace consisting of a graphite susceptor in an insulated, gas-tight, chamber. The furnace was purged with argon to prevent oxidation of the graphite susceptor. Power was supplied by an Ambrell Ekoheat 15/100 radio frequency power supply. Temperature was controlled by a Eurotherm 2416 temperature controller with a Type B thermocouple. Two sample types were used to study matte–refractory interactions: ➤ The first design consisted of crucibles and lids made from the supplied magnesia-chrome brick. Matte and slag were placed into a cavity drilled into a core cut from the refractory brick (Figure 1). It was intended to investigate the interaction between the matte and the refractory brick crucible. In this design the refractory crucible was in direct contact with the graphite susceptor
Figure 1—Crucible design 1 (left) and 2 (right)
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Figure 2—Backscattered electron image of the microstructure of the as-received magnesia-chrome brick
magnesium oxide. The penetration that took place was minimal but uniform throughout the refractory. This experimental set-up presented numerous difficulties as there was no matte left for chemical analysis due to the complete penetration. This design also allowed the refractory to be in direct contact with the graphite susceptor of the induction furnace, thereby exposing it to a very reducing atmosphere. These reducing conditions resulted in the formation of an iron-nickel alloy within the refractory closest to the graphite susceptor. It was subsequently decided to change the design to a set-up in which a small piece of refractory, surrounded with a large amount of matte, was placed in a large alumina crucible (Design 2). The crucible was closed with an alumina lid to limit the sulphur losses from the matte. No slag was used in this design as the slag tended to react with the alumina crucible. Alloy formation in the matte was not observed when Design 2 was used.
Design 2 – alumina crucible containing matte and refractory sample Observed matte-refractory interaction is discussed in terms of matte penetration into the open porosity, matte penetration into the fused aggregate grains, and matte-refractory interaction at the interface.
Matte penetration into open porosity
Figure 3—Backscattered electron image of matte penetration (white phase) into the magnesia-chrome crucible at 1300°C
All the examined refractory samples from Design 2 were completely penetrated by matte, similar to samples from Design 1. A penetration profile could not be distinguished as the refractory samples had a ‘wicking’ effect on the surrounding matte. All the samples had roughly the same outward appearance (Figure 5). The extent of penetration as a function of temperature can be seen in Figure 6. The porosity of the unreacted brick (A) became completely filled with matte as the refractory piece came in contact with the matte at the different temperatures. The large fused grains began disintegrating as the temperature was raised and finally the matte even penetrated into the fused grains themselves.
Matte penetration into the fused aggregate grains Examination of the aggregate grains at higher magnifications gave important insight into the reactions taking place at
Figure 4—Backscattered electron image of matte penetration into the magnesia-chrome crucible at 1300°C (high magnification)
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Figure 5—Refractory sample tested at 1400˚C VOLUME 114
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The white phase is a penetrated Fe-rich sulphide phase, while the very light grey phase represents a Cu-Fe-Ni-based sulphide (Figure 4). The light grey phase in Figure 4 is the typical chromite spinel that is found in the refractory, with the dark grey phase surrounding it consisting of mainly
Wear of magnesia-chrome refractory bricks as a function of matte temperature Matte-refractory interaction at the refractory-matte Interface It was important to carry out a detailed analysis of the interface between the refractory and the surrounding matte, as this would indicate how the actual furnace lining would react when brought in contact with the superheated matte. Tests at 1300˚C and 1400˚C again resulted in matte penetration only.
Figure 6—Reflected light micrographs of matte penetration into the magnesia-chrome refractory as a function of temperature. A: Unreacted brick, B: 1300˚C, C: 1400˚C, D: 1500˚C, E: 1600˚C, F: 1700˚C)
increasing temperatures. Except for penetration, the tests at 1300°C and 1400°C showed no clear signs of chemical interaction. It was only at temperatures of 1500°C and higher that signs of interaction were noticed. This agrees with FactSAGE® predictions. At 1500°C and higher, spinels migrated out of the magnesia grains and collected on the grain boundaries. This left large areas of the magnesia grains exposed as they became depleted in spinels. Furthermore, matte started to penetrate into the fused grains and disintegrated them (Figure 7). EDS analysis of this penetrating matte indicated that it locally contained up to 0.7% Cr, compared to 0.3% Cr in the bulk matte. Disintegration of the fused grains was more pronounced in the refractory samples that were reacted at 1600°C and 1700°C (Figures 8 and 9). The spinel crystals migrated out of the magnesia grains, while the matte completely disintegrated the large fused grains. At 1600°C the penetrating matte had a Cr content of 0.5% compared to 0.3% in the bulk matte, which is not a significant increase. At 1750˚C (Figure 10) the penetrating matte contained 0.8% Cr compared with a bulk matte concentration of 0.3%. A sudden increase in porosity was also observed (Figure 10), together with a large amount of sulphur gas evolution during the experiment. As the experimental temperature increased the spinel crystals increasingly migrated out of the magnesia grains and grouped together as the magnesia grains disintegrated. This destruction of the fused grains would make the bricks more vulnerable to the effects of thermal shock, as without the finely dispersed spinel crystals the bricks would be more susceptible to crack propagation.
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Figure 7—Backscattered electron image of the microstructure of the magnesia-chrome brick reacted at 1500˚C
Figure 8—Backscattered electron image of the microstructure of the magnesia-chrome brick reacted at 1600˚C
Figure 9—Backscattered electron image of the microstructure of the magnesia-chrome brick reacted at 1700˚C The Journal of The Southern African Institute of Mining and Metallurgy
Wear of magnesia-chrome refractory bricks as a function of matte temperature the temperature increased the matte also started to penetrate the fused aggregate grains, and disintegrated them. CrS was not observed as a separate phase. It was noticed, however, that the chromium content of the matte inside the refractory sample was slightly higher than that of the bulk
Figure 10—Backscattered electron image of the microstructure of the magnesia-chrome brick reacted at 1750˚C
Figure 12—Backscattered electron image of the matte-refractory interface at 1600˚C (dark grey phase in the light grey matte is Mg-rich silicate crystals)
Figure 11—Backscattered electron image of the matte-refractory interface at 1500˚C
At temperatures of 1500ºC and higher a magnesium-rich silicate phase could be observed to form a boundary layer at the refractory-matte interface (Figures 11 to 13). This boundary phase could have been established through the predicted formation of MgO that reacted with the silica in either the matte or refractory, as both contained small amounts of silica. On either side of the boundary, FeO crystals started to form, together with spinel crystals. The stoichiometry of these crystals started to change from M3O4 to MO. Some large spinel crystals could also be observed at the refractory-matte interface, as well as magnesia grains with finely dispersed exsolved spinel crystals. At temperatures of 1700°C and 1750°C more FeO crystals started to form at the interface, and the stoichiometry of the magnesium-rich silicate phase was calculated to be (Mg,Fe)2SiO4 (Figures 13 and 14). The extent of spinel segregation out of the magnesia grains close to the interface was also higher at these temperatures. At 1750°C the spinels were completely removed from the fused magnesia grains close to the refractory-matte interface (Figure 14).
Figure 13—Backscattered electron image of the matte-refractory interface at 1700˚C
Conclusions
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The observed amount of matte penetration into the magnesia-chrome brick samples was substantial. All the refractory samples were completely penetrated by matte. As
Wear of magnesia-chrome refractory bricks as a function of matte temperature matte. At temperatures of 1500˚C and higher, definite signs of MgO and FeO formation were found, as well as the gradual disintegration of the grains inside the refractory at the refractory-matte interface. The formation of a magnesium-rich silicate boundary phase at the matte-refractory interface at temperatures above 1500ºC was unexpected and needs further investigation. Results obtained in this study clearly indicate that the reactions that took place between primary PGM matte and the magnesia-chrome refractory material are more complex than those predicted through FactSAGE® modelling.
Southern African Pyrometallurgy, Cradle of Humankind, South Africa, 6-9 March 2011. Jones, R. and den Hoed, P. (eds). Southern African Institute of Mining and Metallurgy, Johannesburg. pp. 295–308. JACOBS, M. 2006. Process discription and abbreviated history of Anglo Platinum’s Waterval Smelter. Southern African Pyrometallurgy, Cradle of Humankind, South Africa, 5-8 March 2006. Jones, R. (ed.). Southern African Institute of Mining and Metallurgy, Johannesburg. pp. 17–28. JONES, R. 1999. Platinum smelting in South Africa. South African Journal of Science, vol. 95. pp. 529–539. JONES, R. 2005. An overview of Southern African PGM smelting. Mintek,
Acknowledgements
Randburg, South Africa.
Technical and financial support from Anglo American Platinum is gratefully acknowledged. This work is based on research supported in part by the National Research Foundation of South Africa (Grant number TP1208219517).
NELL, J. 2004. Melting of platinum group metal concentrates in South Africa. Journal of the South African Institution of Mining and Metallurgy, vol. 104, no. 7. pp. 423–429. SNYDERS, C., EKSTEEN, J., and MOSHOKWA, A. 2006.. The Polokwane smelter
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R.T. 2007. World nonferrous smelter survey part IV: nickel: sulfide. Journal of Metals, April 2007. pp. 58–72.
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The SAIMM Library The library has been indexed and sorted. Although it is not as big as we would like it to be, we have a fair number of reference books and a certain amount of resource material. Access to the library and the control of the borrowing process is in the hands of Kea Shumba. The titles are available on the website at the following link: http://www.saimm.co.za/saimmlibrary?task=showCategory&catid=32 Sam Moolla, Manager: SAIMM Secretariat
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The Journal of The Southern African Institute of Mining and Metallurgy
Comparing the extent of the dissolution of copper-cobalt ores from the DRC Region by S. Stuurman*, S. Ndlovu*, and V. Sibanda* Paper written on project work carried out in partial fulfilment of BSc Engineering (Metallurgical and Materials Engineering)
Inorganic acids such as sulphuric acid have found use together with certain reducing agents in leaching of copper-cobalt oxide ores. These reagents are not ideal due to the adverse effect the inorganic acids generally have on the environment and the high costs of the reducing agents. In this study a copper-cobalt oxide ore from the Central African Copperbelt was leached in two different environments; sulphuric acid in conjunction with hydrogen peroxide as a reducing agent and tartaric acid. The effects of acid concentration, reducing agent concentration, and temperature were independently determined for both leaching environments. The sulphuric acid concentration was varied between 0.4 M and 1.2 M and the concentration of hydrogen peroxide between 4.0 M and 6.5 M, while the tartaric acid concentration was varied between 0.15 M and 0.35 M. The temperature was varied between 20°C and 50°C. The results showed that the extraction of both copper and cobalt increased with sulphuric acid concentration, reaching a peak at approximately 0.8 M and then decreasing at higher acid concentrations. A similar increase and decrease in metal extraction was observed when the reducing agent was increased. In leaching with tartaric acid, the extraction of cobalt was much higher than that of copper, although extraction of both metals increased with acid concentration. Additions of small amounts of hydrogen peroxide were found to increase cobalt extraction in tartaric acid but had a minimal effect on copper. An increase in the solution temperature had a significant effect in the organic acid environment, with the effect on cobalt extraction being much more pronounced than on copper. Keywords leaching, copper, cobalt, sulphuric acid, hydrogen peroxide, tartaric acid, reducing agent.
Introduction Copper and cobalt are commodities of great economic value. Copper derives its value from its outstanding properties which include high electrical and thermal conductivity, malleability, and toughness. Copper is widely used in electrical cabling, piping, and the construction industry. Cobalt owes its value to the fact that it can maintain its ferromagnetism up to temperatures higher than any other metal, and to its superior catalytic qualities. Cobalt is widely used in the manufacture of magnets, catalysts, and batteries. It is also used as a pigment in paints and as an alloying element in steel production. The Journal of The Southern African Institute of Mining and Metallurgy
* School of Chemical and Metallurgical Engineering, University of the Witwatersrand, Johannesburg, South Africa. © The Southern African Institute of Mining and Metallurgy, 2014. ISSN 2225-6253. Paper received Mar. 2014. VOLUME 114
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Synopsis
The Central African Copperbelt hosts 40% of the world’s cobalt reserves and 10% of the world’s copper reserves. The two countries that lie on the Copperbelt, the Democratic Republic of Congo (DRC) and Zambia, produce about half of the world’s copper-cobalt ore (Crundwell et al., 2011). The DRC is said to be the biggest miner of these ores, although the largest cobalt refinery is in China (Miller, 2009). Zambia and the DRC have a very important role to play in supplying global copper and cobalt requirements. Escalating global demand for copper and cobalt has forced companies to increasingly exploit lower-grade ores. However, most of these low-grade ores are processed at very high costs, since large volumes have to be processed. The methods applied for the extraction of copper and cobalt from the copper-cobalt oxide ores of the Central African Copperbelt include (i) heap leaching of the low-grade ores, (ii) upgrading of the ores by flotation prior to leaching, and (iii) direct ore leaching using sulphuric acid. All these methods have inherent problems. For instance, heap leaching often has low rates of recovery, the inclusion of a flotation stage prior to leaching is costly, and direct ore leaching consumes large volumes of acid. The main cobalt-bearing mineral in the copper-cobalt oxide ore deposits of the DRC is heterogenite (CoO.2Co2O3.6H2O). Copper typically occurs as chrysocolla (CuOSiO2.2H2O) and malachite [CuCO3.Cu(OH)2] (Crundwell et al., 2011). Minor amount of copper silicates such as dioptase (CuSiO3.H2O), katangite (CuS13O9.nH2O), and carbonates such as azurite (Cu3(OH)2(CO3)2 also occur (Prasad, 1989). The main oxides are listed in Table I.
Comparing the extent of the dissolution of copper-cobalt ores from the DRC Region Table I
Primary copper and cobalt oxides found in copper-cobalt ores of the Democratic Republilc of Congo (Prasad, 1989) Oxides of copper
Oxides of Cobalt
Malachite
Cu2(OH)2CO3
Heterogenite
(CO2O3.CuO)H2O
Cuprite
Cu2O
Kolwezite
(Cu, Co)2(CO3)(OH)2
Libethenite
Cu2(OH)PO4
Stainierite
Co2O3.H2O
Pseudomalachite
Cu2(OH)4PO4H2O
Amorphous heterogenite
CoO.2Co2O3.6H2O
Tenorite
CuO
In the Katanga province of the DRC, cobalt is commonly produced from heterogenite (CoO.2Co2O3.6H2O) as a byproduct of copper production. Cobalt in heterogenite occurs in both the 2+ and 3+ oxidation states. In contrast to copper oxide minerals, which readly dissolve in sulphuric acid solution, cobalt is difficult to leach from heterogenite. This is because when heterogenite is leached in an inorganic acid such as sulphuric acid, Co2+ easily goes into solution, but Co3+ does not dissolve: [1] The insoluble Co3+ becomes soluble only after reduction to Co2+. Therefore the hydrometallurgical dissolution of Co3+ can take place only in the presence of a reducing agent. The reducing agents commonly used in the process include ferrous ions (present in the leach solution as a result of leaching of iron minerals in the ore and leaching of iron scrap), sodium metabisulphite (Na2S2O5), and hydrogen peroxide (Apua and Mulaba-Bafubiandi, 2011; Lydall and Auchterlonie, 2011; Seoa et al., 2013). The consumption of these reducing agents is generally high and these make up 47% of the total operation which makes the production of copper and cobalt very expensive (Mulaba-Bafubiandi et al., 2007). The use of sulphuric acid as the main leaching agent also poses a threat to the environment. Sulphuric acid waste leach solution, because it is non-biodegradable and toxic, is stored in on-site holding ponds, which are securely lined. Even though the holding ponds are considered environmentally safe there have been reports of spills impacting groundwater even miles away (Freeman, 2005). Therefore, an economical method of extracting the minerals that uses a more environmentally friendly leaching reagent is required in order for the processing of these low-grade ores to be profitable. Since inorganic acids have not performed well in economic terms or in meeting the standards for ‘green’ chemistry and environmental impact, this has prompted the need to consider organic acids as lixiviants. To date, organic acids have not been used widely as leachants due to their reported low leaching efficiencies. They are also very expensive and are, as a result, most unlikely to be used on low-grade ores in conventional processing routes. However, they are attractive due to the fact that they have been found to cause less harm to the environment as they are biodegradable. In addition, organic acids are also recyclable
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(Gharabaghi et al., 2010). Organic acids extract metals by forming complexes. The more stable the complex the higher the extraction rate, the smaller the ionic radius of the target metal the more stable the complex formed, and the higher the oxidation state of the metal the more stable the metal complex (Weisstein, 2011). Currently there are no documented studies on the leaching of copper-cobalt oxide ores from the Central African Copperbelt by organic acids. In the present work, preliminary studies have been conducted to establish the feasibility of leaching using an organic acid. The leaching behaviour of a copper-cobalt oxide ore in an organic acid environment (tartaric acid) was compared with that in an inorganic environment (sulphuric acid) fortified with a reducing agent (hydrogen peroxide). Tartaric acid has been observed to chelate metal ions and is relatively cheap compared to other organic acids. This choice was further reinforced by studies comparing the leaching rates of heavy metals from contaminated soils and spent batteries using different organic acids, which showed that tartaric acid and citric acid could remove heavy metals from contaminated soils and waste material more efficiently and rapidly than all other potential organic extractants (Wasay et al., 2001; Li et al., 2010).
Materials and methods Ore sample The ore used in the test work is a copper-cobalt oxide ore from the Katanga Province in the DRC. The ore was crushed and milled to 80% –150 μm. The composition of the ore is given in Table II.
Reagents The main reagents included analytical grade sulphuric acid (98%) as a leaching reagent, analytical grade hydrogen
Table II
Chemical composition of the Cu-Co ore sample used in the test work Element Wt%
Cu
Co
Zn
Ni
Fe
Mn
Mg
SiO2
4.53
0.3
0.029
0.003
1.4
0.22
2.48
84.5
The Journal of The Southern African Institute of Mining and Metallurgy
Comparing the extent of the dissolution of copper-cobalt ores from the DRC Region peroxide (30%) as the reducing agent, 100% tartaric acid, and distilled water for dilution. All reagents were sourced from Merck Millipore.
Experimental method Inorganic acid leaching Dilute solutions of sulphuric acid and hydrogen peroxide were prepared separately. The dilute solutions were mixed in beakers accordingly to obtain the desired pH. The beakers were then placed in a shaking water bath that was set at the desired temperature and the solutions were given time to reach a steady temperature. The pH of the solutions was monitored using a calibrated pH meter and the temperature was measured using a thermometer. Once the solution pH and temperature of the solution were stable, the ore sample was introduced into the solution in the beakers. The pulp in the beaker was continuously stirred for 3 hours to facilitate leaching. Table III summarizes the test conditions for the inorganic acid leaching tests. The experiments were all undertaken over a 3 hour period.
Organic acid leaching Dilute acid solutions of different concentrations were prepared by dissolving tartaric acid powder in distilled water. The pH of the solutions was measured. The solutions were poured into beakers and the beakers were placed in a shaking water bath to achieve the desired temperature. Once the leaching solution in the beaker reached a steady temperature, the ore sample was introduced. The pulp was left to leach for 24 hours while being continuously stirred. A sample for analysis was taken after the first 3 hours. Table IV summarizes the test conditions for the organic acid leaching tests. The experiments were all undertaken over a 3 hour period.
Leachate analysis Leachate samples were analysed to determine the concen-
tration of copper and cobalt after each experiment. In both sets of experiments, the leached pulp was left to settle in order to allow for solid/liquid separation .. to occur. The pulp was then filtered using filter paper, a Bu–chner funnel, and a filter flask to obtain the leach solution. Samples of the filtered solution were poured into sample bottles and sent for copper and cobalt analysis by inductively coupled plasma mass spectrometry (ICP-MS) using a Perkin Elmer Nexlon 300D instrument.
Results and discussion Inorganic acid leaching Effect of sulphuric acid concentration Figure 1 shows the extraction of the metals as a function of sulphuric acid concentration. The results show similar trends for the extraction of both copper and cobalt. At 0.4 M acid concentration, cobalt and copper extractions were 34.0% and 35.0% respectively. Extraction increased to a maximum of 97.4% (Cu) and 78.2% (Co) at 0.8 M sulphuric acid concentration, and then declined with further increases in acid concentration to 78.3% for copper and 63.3% for cobalt at 1.2 M. One would expect that with an increasing acid concentration the dissolution of the copper and cobalt would
Figure 1—Effect of sulphuric acid concentration on copper and cobalt extraction (temperature 25°C, hydrogen peroxide concentration 3M, leaching time 3 hours
Table III
Test conditions for inorganic acid leacing in a reducing environment Test
Temperature (°C)
Effect of sulphuric acid concentration Effect of hydrogen peroxide (reducing agent) concentration Effect of temperature
Sulphuric acid concentration (M) Hydrogen peroxide concentration (M)
25 25 20, 30, 40
0.4, 0.8, 1.2 0.4 0.8
3.0 4.0, 5.5, 6.5 4.0
Temperature (°C)
Tartaric acid concentration (M)
Hydrogen peroxide concentration (M)
25 25 20, 30, 40, 50
0.15, 0.25, 0.35 0.35 0.35
0.0 3.0, 4.5, 5.5, 6.5 0.0
Table IV
Test Effect of tartaric acid concentration Effect of hydrogen peroxide (reducing agent) concentration Effect of temperature
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Comparing the extent of the dissolution of copper-cobalt ores from the DRC Region increase, as there are more H+ ions available. However, this is not the case. The Eh-pH diagrams for copper and cobalt in Figure 2 and Figure 3 show that copper and cobalt ionization is favoured at relatively low pH levels. The diagrams indicate that in order to solubilize copper as Cu2+, the copper oxide minerals have to be leached in highly acidic conditions (pH <1). At pH conditions >1, a solid CuOFeO2 is formed (Figure 2). Similarly, cobalt is best leached at pH <1. At pH levels >1, a solid CoO.Fe2O3 is formed, as seen in Figure 3, which has a passivating effect on the leaching of cobalt. Thus as the pH increases beyond about 2 there is a possibility of formation of solid compounds of copper (CuFeO2) and cobalt (CoO.Fe2O3). These solids have a passivating effect on the leaching reactions, resulting in a decrease in the dissolution of both copper and cobalt. Solutions >1.0 M in concentration had a starting pH greater than 1. Addition of ore may have increased the pH to levels that favour passivation. The shift from increasing extraction to decreasing extraction (Figure 1) may also be due to a change in the ratecontrolling step. At acid concentrations between 0.4 M and 0.8 M, the reaction kinetics may be controlled by the mass transfer of acid from the bulk solution to the particles. At concentrations from 0.8 M to 1.2 M, the controlling
Figure 2—Eh-pH diagram of the Cu-O-Fe system at 25°C (Schlesinger et al., 2011)
Figure 3—Eh-pH diagram of the Co-O-Fe system at 25°C (Schlesinger et al., 2011)
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mechanism may have changed, so that further increases in acid concentration would have no effect in the dissolution of both cobalt and copper ores.
Effect of the concentration of the reducing agent The reaction of Co3+ oxides in the presence of hydrogen peroxide as the reducing agent is expected to proceed by the following reactions: [2]
[3] The effect of a reducing agent on metal recoveries was investigated in preliminary test work that focused on the solution Eh. In the absence of reducing agent the Eh of the solution varied between 600 and 900 mV vs SCE (0.242 V) over the leaching period. However, on addition of about 3.0 M hydrogen peroxide the Eh of the solution decreased drastically to values around 300 to 500 mV vs SCE. According to Figure 1, the lower Eh of the solution subsequently enhances the dissolution of Co (III) phase. The effect of hydrogen peroxide concentration in conjunction with sulphuric acid was then tested. Figure 4 shows the metal extraction as a function of reducing agent concentration. The results show similar trends for both copper and cobalt extraction. Maximum extractions of 95.1% for copper and 79.4% for cobalt were obtained at around 4 M hydrogen peroxide. Extraction of copper and cobalt then decreased to 59.3% and 63.8% respectively at about 6.5 M hydrogen peroxide. It is important to note that copper extraction was higher than that of cobalt at all hydrogen peroxide concentrations except 6.5 M. This possible indicates different reaction mechanisms controlling the Cu and Co extraction processes. Another point is the sharp increase in metal extraction from 3.0 M to 4.0 hydrogen peroxide concentration, which is followed by a decrease in extraction at 5.5 M and above. This clearly indicates that the concentration of reducing agent has a positive effect on metal extraction only up to a certain extent; above this range the reaction rate becomes less dependent on hydrogen peroxide concentration. At reducing agent concentrations above 4 M, the mass transfer of reducing agent from solution to particles may no longer be the rate controlling mechanism.
Figure 4—Effect of hydrogen peroxide concentration on copper and cobalt extraction (temperature 25°C, sulphuric acid concentration 0.4 M, leaching time 3 hours) The Journal of The Southern African Institute of Mining and Metallurgy
Comparing the extent of the dissolution of copper-cobalt ores from the DRC Region Since the optiuum cobalt and copper extractions were obtained with 4.0 M hydrogen peroxide and 0.4 M sulphuric acid, these conditions were used in the temperature tests. Figure 5 shows the metal extractions as a function of temperature. There is a marginal increase in the extraction of both copper and cobalt as the leaching temperature is increased. It is known from the Arrhenius equation that an increase in temperature generally enhances reaction kinetics. Increasing temperature may thus increase the mass transfer kinetics of acid and reducing agent from solution to particles. The reaction at the surface of the particles is also enhanced by increasing temperature, thus leading to increased metal dissolution and higher extraction efficiency. However, the temperature increase does not seem to have as great an effect on the extraction efficiency of copper and cobalt as compared to acid and reducing agent concentration. The temperature effect was therefore capped at 40°C.
Organic acid leaching Effect of tartaric acid concentration Figure 6 shows the extraction of copper and cobalt as a function of tartaric acid concentration. The copper and cobalt extractions are both much lower than those achieved with the inorganic acid (Figure 1), with the difference being more significant for copper than for cobalt. The maximum copper extracted within the range of tartaric acid concentration used was about 5% at 0.35 M, compared to about 90% in sulphuric acid media. The maximum cobalt extracted was about 38% in tartaric acid compared to about 80% in sulphuric acid media. The dissolution of copper and cobalt is governed by the extent to which the two acids dissociate, which is quantified by their respective dissociation constants (pKa). The lower the pKa value, the higher the dissociation rate. Sulphuric acid has a pKa of 1.99, compared with the pKa1 of 2.98 and pKa2 of 4.34 for tartaric acid at 25°C (Murthy, 2008). It is also noticeable from Figure 6 that the maximum leaching ability of tartaric acid was not reached within the selected concentration range. The effect of increased tartaric acid concentrations on cobalt extraction should be investigated in future studies. Cobalt extraction was much higher than that of copper at all acid concentrations tested. This is notably the reverse of what was observed with the sulphuric acid leaching process,
Figure 5—Effect of temperature on copper and cobalt dissolution. Sulphuric acid concentration 0.4 M, hydrogen peroxide concentration 4 M, leaching time 3 hours The Journal of The Southern African Institute of Mining and Metallurgy
where the extraction of copper was significantly higher than that of cobalt Furthermore, the level of copper extraction did not change significantly with increasing acid concentration. These observations indicate that the cobalt mineral has a faster reaction rate with the organic acid than the copper mineral, and that the leaching reactions of the two metal systems may be controlled by two different mechanisms. Extraction of metals from ores by organic acids generally takes place through protonation, chelation, and ligand exchange reactions. Thus the organic acid supplies both protons and metal-complexing organic acid anions, with the protons contributing to proton-promoted mineral dissolution. The major factor with organic acids, however, may be that metal-organic complexes can form at the solid-solution interface, weakening cation-oxygen bonds and catalysing the dissolution reaction. The greater the stability of the metal complex formed; the higher the metal dissolution. The stability of the complex depends on the ionic radius and the oxidation state of the metal ion. A smaller ionic radius and a higher oxidation state both increase the stability of the metal complex formed.. Unlike Co (II) complexes, Co (III) complexes undergo ligand substitution reactions relatively slowly and so tend to be stable to ligand exchange (Kim et al., 1993). This could explain the lower cobalt extraction in the organic acid as compared to sulphuric acid enhanced with a reducing agent. This might indicate that the Co (II) mineral species undergo dissolution through the protonation mechanism. The resulting Co (III) species, however, did not undergo much dissolution due to the slow ligand exchange reactions. In view of the low extractions observed with tartaric acid, the effect of additions of a reducing agent was investigated.
Figure 6—Effect of tartaric acid concentration on copper and cobalt extraction (temperature 25°C, leaching time 3 hours)
Figure 7—Effect of reducing agent concentration on copper and cobalt extraction (tartaric acid concentration 0.35 M, temperature 25°C, leaching time 3 hours) VOLUME 114
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Effect of temperature
Comparing the extent of the dissolution of copper-cobalt ores from the DRC Region Effect of addition of a reducing agent
Conclusions
Figure 7 shows the effect of hydrogen peroxide additions to the organic acid leach solution. The addition of a reducing agent had a significant effect on cobalt extraction, but little effect on copper. As outlined in the preceding section it is possible that the addition of a reducing agent will reduce the Co (III) species to Co (II).Furthermore, the uncomplexed Co (III) ion itself is not stable in water due to the hydrolysis/reduction reaction:
Sulphuric acid in the presence of hydrogen peroxide as a reducing agent was able to extract a large amount of both copper and cobalt. About 95% copper and 80% cobalt extractions were achieved in 0.4 M sulphuric acid and 4 M hydrogen peroxide in 3 hours’ leaching time at 25°C. Extractions of both copper and cobalt increased with an increase in sulphuric acid concentration up to 0.8 M. Addition of hydrogen peroxide to the sulphuric acid leaching solution had a positive effect on both copper and cobalt dissolution up to 4 M hydrogen peroxide. An increase in temperature, however, did not have as an significant effect on the extraction efficiency of copper and cobalt as the reducing agent concentration. With tartaric acid as the lixiviant, about 40% cobalt and 5% copper extraction were realized at 25°C. Thus more copper and cobalt were extracted in the inorganic acid environment than in the organic acid environment. However, the addition of hydrogen peroxide to the tartaric acid leaching solution resulted in an 80% cobalt extraction and about 10% copper extraction under the same temperature conditions. In the tartaric acid leaching environment, the change in temperature had a much more pronounced effect on cobalt extraction than that of copper, with about 60% cobalt and 10% cobalt extracted in the absence of hydrogen peroxide. In addition, changes in temperature had a more significant effect on the extraction of cobalt in organic solutions than in the inorganic environment. The results obtained in these two leaching environments indicate the potential of tartaric acid to extract cobalt, rather than copper, from the copper-cobalt ores. It is recommended that further investigations be carried out with higher concentrations of tartaric acid, since in the current test work the concentration of tartaric acid was too low to achieve maximum leaching capability. Tests involving other commonly used organic acids such as citric and oxalic acid would also add value to this research area.
[4] The generated Co (II) ions undergo ligand substitution much faster, and as a result cobalt dissolution will be greatly enhanced.
Effect of temperature Figure 8 shows the effect of temperature on the extraction of cobalt and copper in tartaric acid in the absence of hydrogen peroxide. Tests were carried out from 20°C to 50°C over a 3 hour leaching period. As with all chemical reactions, the extraction of both cobalt and copper increases with increasing temperature. Figure 8 shows that in general, an increase in temperature has a much more positive effect on the extraction of cobalt than that of copper. The highest cobalt extraction (62%) was recorded at the highest temperature used (50°C), while only 10% copper extraction was recorded at the same temperature. According to Shabani et al. (2012) the dissolution of cobalt is highly dependent on temperature, with the effect being more noticeable in the presence of tartaric acid than in sulphuric acid (Figure 5). Organic acids are weak acids and their metal dissolution abilities are affected by the extent of dissociation. The more the acid dissociates, the greater its ability to solubilize metals. An increase in temperature results in increased dissociation of tartaric acid, increasing the number of hydrogen and ligand ions in the acid and thus enhancing the cobalt extraction. One other noteworthy aspect is that the temperature tests were carried out in the absence of a reducing agent. It is possible therefore that an increase in temperature beyond 50°C could result in the extraction efficiency for cobalt reaching similar levels to those observed at lower temperatures in the presence of hydrogen peroxide. This aspect should be investigated further in future work, as it could have an impact on the potential applicability and economics of the process.
Acknowledgements The authors would like to thank Gecamines in DRC for the provision of the ore samples used in this study and the School of Chemical and Metallurgical Engineering at the University of the Witwatersrand for technical assistance.
References APUA, M. and MULABA-BAFUBIANDI, A. 2011. Dissolution of oxidised Co–Cu ores using hydrochloric acid in the presence of ferrous chloride. International Journal of Chemical and Biological Engineering, 28 April. pp. 47–51. CRUNDWELL, F.K., MOATS, M., RAMACHANDRAN, V., ROBINSON, T., and DAVENPORT, W.G. 2011. Extractive Metallurgy of Nickel, Cobalt and Platinum Group Metals. Elsevier, Amsterdam. FREEMAN, N.F. 2005. ADEQ. http://www.savethesantacruzaquifer. info/index.htm[Accessed 13 October 2013]. GHARABAGHI, M., IRANNAJAD, M., and NOAPARAST, M. 2010. A review of the beneficiation of calcareous phosphate ores using organic acid leaching. Hydrometallurgy, vol. 103. pp. 96–107 KIM, J.H., BRITTEN, J., and CHIN, J. 1993. Kinetics and mechanism of Cobalt(III)
Figure 8—Effect of temperature on copper and cobalt extraction (tartaric acid concentration 0.35 M, leaching time 3 hours)
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complex catalysed hydration of nitriles. Journal of the American Chemical Society, vol. 115. pp. 3618–3622. The Journal of The Southern African Institute of Mining and Metallurgy
Comparing the extent of the dissolution of copper-cobalt ores from the DRC Region LI, L., GE, J., WU, F., CHEN, R., CHEN, S., and WU, B. 2010. Recovery of cobalt and lithium from spent lithium ion batteries using organic citric acid as leachant. Journal of Hazardous Materials, vol. 176. pp. 288–293. LYDALL, M.I. and AUCHTERLONIE, A. 2011. The Democratic Republic of Congo and Zambia: a growing global ‘hotspot’ for copper-cobalt mineral investment and exploitation. 6th Southern Africa Base Metals Conference, Phalaborwa, 18–20 July 2011. Southern African Institute of Mining and Metallurgy, Johannesburg. pp. 25–38.
PRASAD, M.S. 1989. Production of copper and cobalt at Gecamines, Zaire. Minerals Engineering, vol. 2, no. 4. pp. 521–541. SHABANI, M.A., IRANNAJAD, M., and AZADMEHR, A.R. 2012. Investigation on leaching of malachite by citric acid. International Journal of Minerals, Metallurgy and Materials, vol. 19, no. 9. pp. 782–786. SCHLESINGER, M.E., KING, M.J., SOLE, K.C., and DAVENPORT, W.G. 2011. Extractive Metallurgy of Copper. 5th edn. Elsevier, Amsterdam. pp. 282-283 SEOA, S.Y., CHOIA, W.S., KIMA, M.J., and TRANA, T. 2013. Leaching of a Cu-Co
MILLER, G. 2009. Design of copper-cobalt hydrometallurgical circuits. ALTA Nickel– Cobalt Conference (ALTA 2009), Perth, WA. MULABA-BAFUBIANDI, A.F., NDALAMO, J., and MAMBA, B. 2007. Microwaveassisted sulphur dioxide flushed acid leaching of mixed cobalt copper oxidised ores. 4th Southern African Base Metals Conference, Swakopmund, Namibia, 23-25 July 2007. Southern African Institute of Mining and Metallurgy, Johannesburg. pp. 9–27. MURTHY, C.P., ALI, S.F.M., DUBEY, P.K., and ASHOK, D. 2008. University Chemistry. vol. 2. New Age International Publishers, New Delhi, India.
ore from Congo using sulphuric acid-hydrogen peroxide leachants. Journal of Mining and Metallurgy, Section B: Metallurgy, vol. 49, no. 1. pp. 1–7. WASAY, S.A., BARRINGTON, S., and TOKUNAGA S. 2001. Organic acids for the insitu remediation of soils polluted by heavy metals: soil flushing in columns. Water Air Soil Pollution, vol. 127. pp. 301–314. WEISSTEIN, E. Complex ion stability. Weisstein’s World of Chemistry. http://scienceworld.wolfram.com/chemistry/ComplexIonStability.html [Accessed 23 October 2013].
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SAIMM 120th Anniversary
he Southern African Institute of Mining and Metallurgy (SAIMM) has redesigned our logo to coincide with our 120th Anniversary. This logo is more aligned with the changes over the last two decades, while maintaining the professionalism that the SAIMM is renowned for. We have also emphasized the fact that we are 120 years old, and have continued to maintain our technical excellence with regard to our Journal and the events that we organize. To add to these achievements we continue to increase our membership.
T
The Parts of an Achievement of Arms and their Significance The arms under consideration comprise separate parts, viz., Shield, Helm, Mantling, Crest, Supporters, Compartment, and Motto.
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The Shield: This is in blue divided by a golden chevron, to represent the major sections of the industry. The flaming crucibles in the upper section represent Metallurgy and the crossed pick and shovel in the lower section represent Mining. The Helm: This is an Esquireʼs Helmet, which is the customary type of use for the arms of corporate bodies. The Wreath and the Mantling: These are always in the two main ʻcoloursʼ of the shield, in this case gold as a metal and blue as the colour. The mantling was originally a short cloak draped from the helmet as a protection against the sun, and the wreath helped to hold the crest in place. The Crest: This served as an additional mark of distinction. In this case the demi-lion represents strength and holds the national flower of South Africa in his left Claw. The Supporters: In this case heraldic beasts have been chosen, symbols of these ancient professions, the black lion representing mining and the golden dragon representing metallurgy. The ʻdifferentʼ marks on their shoulders are carried over from the shield of the Chemical, Mining and Metallurgical Society, and their colours and the diamonds in their collars are intended to represent the main fields of mining in South Africa, namely gold, coal, and diamonds. The Compartments: This is, appropriately, an outcrop of rock. The Motto: ʻCapaci Occasioʼ has been taken over from the Instituteʼs predecessor, the Chemical, Mining and Metallurgical Society, with the exhortation, ʻto the capable the opportunityʼ.
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Forthcoming SAIMM events...
IP PONSORSH EXHIBITS/S ng to sponsor ishi Companies w ese t at any of th and/or exhibi contact the events should rdinator -o conference co ssible as soon as po
SAIMM DIARY or the past 120 years, the Southern African Institute of Mining and Metallurgy, has promoted technical excellence in the minerals industry. We strive to continuously stay at the cutting edge of new developments in the mining and metallurgy industry. The SAIMM acts as the corporate voice for the mining and metallurgy industry in the South African economy. We actively encourage contact and networking between members and the strengthening of ties. The SAIMM offers a variety of conferences that are designed to bring you technical knowledge and information of interest for the good of the industry. Here is a glimpse of the events we have lined up for 2013. Visit our website for more information.
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2014 ◆ SEMINAR Society of Mining Professors— A Southern African Silver Anniversary 26–30 June 2014, The Maslow Hotel, Sandton, Gauteng ◆ SCHOOL Mine Planning School 15–16 July 2014, Mine Design Lab, Chamber of Mines Building, The University of the Witwatersrand ◆ CONFERENCE Pyrometallurgical Modelling Principles and Practices 4–5 August 2014, Misty Hills Country Hotel and Conference Centre, Cradle of Humankind ◆ CONFERENCE MinPROC 2014 6–8 August 2014, Cape Town ◆ CONFERENCE MineSafe Conference 2014 20–21 August 2014, Conference 22 August 2014, Industry day Emperors Palace, Hotel Casino Convention Resort, Johannesburg ◆ SCHOOL 3rd Mineral Project Valuation School 9–11 September 2014, Mine Design Lab, Chamber of Mines Building, The University of the Witwatersrand ◆ CONFERENCE Surface Mining 2014 16–17 September 2014, The Black Eagle Room, Nasrec Expo Centre ◆ SCHOOL Grade control and reconciliation 23–24 September 2014, Moba Hotel, Kitwe, Zambia
◆ CONFERENCE
For further information contact: Conferencing, SAIMM P O Box 61127, Marshalltown 2107 Tel: (011) 834-1273/7 Fax: (011) 833-8156 or (011) 838-5923 E-mail: raymond@saimm.co.za
SHAPE: 1st International Conference on Solids Handling and Process Engineering 29–30 September 2014, University of Pretoria, South Africa
◆ CONFERENCE 6th International Platinum Conference 20–234 October 2014, Sun City, South Africa
Website: http://www.saimm.co.za
INTERNATIONAL ACTIVITIES 2014
19–20 May 2014 — Drilling and Blasting Swakopmund Hotel & Entertainment Centre, Swakopmund, Namibia Contact: Raymond van der Berg Tel: +27 11 834-1273/7, Fax: +27 11 838-5923/833-8156 E-mail: raymond@saimm.co.za, Website: http://www.saimm.co.za 19–23 May 2014 — Fundamentals of Process Safety Management (PSM) Johannesburg, South Africa Contact: RDC Prior Tel: +27 (0) 825540010, E-mail: r.prior@mweb.co.za 24–31 May 2014 — ALTA 2014 Nickel-Cobalt-Copper, Uranium-REE and Gold-Precious Metals Conference & Exhibition Perth, Western Australia Contact: Allison Taylor E-Mail: allisontaylor@altamet.com.au, Tel: +61 (0)411 692-442 Website: http://www.altamet.com.au/conferences/alta-2013/ 27–29 May 2014 — Furnace Tapping Conference 2014 Misty Hills Country Hotel and Conference Centre, Cradle of Humankind Contact: Raymond van der Berg Tel: +27 11 834-1273/7, Fax: +27 11 838-5923/833-8156 E-mail: raymond@saimm.co.za, Website: http://www.saimm.co.za
6–8 August 2014 — MinPROC 2014 Lord Charles Hotel, Somerset West, Cape Town 20–22 August 2014 — MineSafe Conference 2014 Technical Conference and Industry day 20–21 August 2014: Conference 22 August 2014: Industry day Emperors Palace, Hotel Casino Convention Resort, Johannesburg Contact: Raymond van der Berg Tel: +27 11 834-1273/7, Fax: +27 11 838-5923/833-8156 E-mail: raymond@saimm.co.za, Website: http://www.saimm.co.za 18–19 November 2014 — Third International Engineering Materials and Metallurgy Conference and Exhibition (iMat 2014) Shahid Beheshti International Conference Center, Tehran, Iran Contact: Kourosh Hamidi E-mail: info@imatconf.com 9–11 September 2014 — 3rd Mineral Project Valuation School Mine Design Lab, Chamber of Mines Building, The University of the Witwatersrand Contact: Camielah Jardine, Tel: +27 11 834-1273/7, Fax: +27 11 838-5923/833-8156 E-mail: camielah@saimm.co.za, Website: http://www.saimm.co.za 16–17 September 2014 — Surface Mining 2014 The Black Eagle Room, Nasrec Expo Centre Contact: Camielah Jardine, Tel: +27 11 834-1273/7, Fax: +27 11 838-5923/833-8156, E-mail: camielah@saimm.co.za, Website: http://www.saimm.co.za
11–12, June, 2014 — AIMS 2014: 6th International Symposium ‘High Performance Mining’ Aachen, Germany Contact: Sandra Zimmermann Tel: +49-(0)241-80 95673, Fax: +49-(0)241-80 92272 E-Mail: zimmermann@bbk1.rwth-aachen.de Website: http://www.aims.rwth-aachen.de
2015
26–30 June 2014 — Society of Mining Professors A Southern African Silver Anniversary The Maslow Hotel, Sandton, Gauteng, South Africa Contact: Raymond van der Berg Tel: +27 11 834-1273/7, Fax: +27 11 838-5923/833-8156 E-mail: raymond@saimm.co.za, Website: http://www.saimm.co.za
14–17 June 2015 — European Metallurgical Conference Dusseldorf, Germany Website: http://www.emc.gdmb.de
15–16 July 2014 — Mine Planning School Mine Design Lab, Chamber of Mines Building, The University of the Witwatersrand Contact: Camielah Jardine Tel: +27 11 834-1273/7, Fax: +27 11 838-5923/833-8156 E-mail: camielah@saimm.co.za, Website: http://www.saimm.co.za 4–5 August 2014 — Pyrometallurgical Modelling Principles and Practices Emperors Palace Hotel Casino Convention Resort, Johannesburg Contact: Camielah Jardine Tel: +27 11 834-1273/7, Fax: +27 11 838-5923/833-8156 E-mail: camielah@saimm.co.za, Website: http://www.saimm.co.za The Journal of The Southern African Institute of Mining and Metallurgy
March 2015 — PACRIM 2015 Hong Kong, China Tel: +27 11 834-1273/7, Fax: +27 11 838-5923/833-8156, E-mail: CBenn@ausimm.com.au Website: http://www.pacrim2015.ausimm.com.au
14–17 June 2015 — Lead Zinc Symposium 2015 Dusseldorf, Germany Website: http://www.pb-zn.gdmb.de 16–20 June 2015 — International Trade Fair for Metallurgical Technology 2015 Dusseldorf, Germany Website: http://www.metec-tradefair.com 5–9 October 2015 — MPES 2015: 23rd International Symposium on Mine Planning & Equipment Selection Sandton Convention Centre, Johannesburg, South Africa Contact: Raj Singhai E-mail: singhal@shaw.ca or Contact: Raymond van der Berg Tel: +27 11 834-1273/7, Fax: +27 11 838-5923/833-8156 E-mail: raymond@saimm.co.za, Website: http://www.saimm.co.za APRIL 2014
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12–14 May 2014 — 6th South African Rock Engineering Symposium SARES 2014 Creating value through innovative rock engineering Misty Hills Country Hotel and Conference Centre, Cradle of Humankind Contact: Raymond van der Berg Tel: +27 11 834-1273/7, Fax: +27 11 838-5923/833-8156 E-mail: raymond@saimm.co.za, Website: http://www.saimm.co.za
Company Affiliates The following organizations have been admitted to the Institute as Company Affiliates AECOM SA (Pty) Ltd
Elbroc Mining Products (Pty) Ltd
Osborn Engineered Products SA (Pty) Ltd
AEL Mining Services Limited
eThekwini Municipality
Outotec (RSA) (Proprietary) Limited
Air Liquide (Pty) Ltd
Evraz Highveld Steel and Vanadium Limited
PANalytical (Pty) Ltd
AMEC GRD SA
Exxaro Coal (Pty) Ltd
Paterson and Cooke Consulting Engineers
AMIRA International Africa (Pty) Ltd
Exxaro Resources Limited
Paul Wurth International SA
ANDRITZ Delkor(pty) Ltd
Fasken Martineau
Anglo Operations (Pty) Ltd
FLSmidth Minerals (Pty) Ltd (FFE001)
Polysius A Division Of Thyssenkrupp Engineering
Anglogold Ashanti Ltd
Fluor Daniel SA ( Pty) Ltd
Arcus Gibb (Pty) Ltd
Franki Africa (Pty) Ltd-JHB
Atlas Copco Holdings South Africa (Pty) Limited
Fraser Alexander Group
Aurecon South Africa (Pty) Ltd Aveng Mining Shafts and Underground Aveng Moolmans (Pty) Ltd Bafokeng Rasimone Platinum Mine Barloworld Equipment -Mining BASF Holdings SA (Pty) Ltd Bateman Minerals and Metals (Pty) Ltd BCL Limited (BCL001) Becker Mining (Pty) Ltd BedRock Mining Support Pty Ltd
Precious Metals Refiners Rand Refinery Limited Redpath Mining South Africa (Pty) Ltd Rosond (Pty) Ltd
Goba (Pty) Ltd
Royal Bafokeng Platinum
Hall Core Drilling (Pty) Ltd Hatch (Pty) Ltd
Roymec Technologies (Pty) Ltd
Herrenknecht AG
RSV Misym Engineering Service (Pty) Ltd
HPE Hydro Power Equipment (Pty) Ltd
RungePincockMinarco Limited
Impala Platinum Holdings Limited
Rustenburg Platinum Mines Limited
IMS Engineering (Pty) Ltd
SAIEG
JENNMAR South Africa
Salene Mining (Pty) Ltd
Joy Global Inc.(Africa)
Sandvik Mining and Construction Delmas (Pty) Ltd
Leco Africa (Pty) Limited
Bell Equipment Limited
Longyear South Africa (Pty) Ltd
Sandvik Mining and Construction RSA(Pty) Ltd
BHP Billiton Energy Coal SA Ltd
Lonmin Plc
SANIRE
Blue Cube Systems (Pty) Ltd
Ludowici Africa (Pty) Ltd
Sasol Mining (Pty) Ltd
Bluhm Burton Engineering Pty Ltd
Wekaba Engineering (Pty) Ltd
Scanmin Africa (Pty) Ltd
Blyvooruitzicht Gold Mining Company Ltd
Magnetech (Pty) Ltd
Sebilo Resources (Pty) Ltd
BSC Resources Ltd
MAGOTTEAUX (PTY) LTD
SENET (Pty) Ltd
CAE Mining (Pty) Limited
MBE Minerals SA Pty Ltd
Senmin International (Pty) Ltd
Caledonia Mining Corporation
MCC Contracts (Pty) Ltd
Shaft Sinkers (Pty) Limited
CDM Group
MDM Technical Africa (Pty) Ltd
Sibanye Gold Limited
CGG Services SA
Metalock Industrial Services Africa (Pty)Ltd
Smec SA
Chamber of Mines
Metorex Limited
SMS Siemag
Concor Mining
Metso Minerals (South Africa) (Pty) Ltd
SNC Lavalin (Pty) Ltd
Concor Technicrete
Minerals Operations Executive (Pty) Ltd
Sound Mining Solution (Pty) Ltd
Council for Geoscience
MineRP
SRK Consulting SA (Pty) Ltd
CSIR Natural Resources and the Environment
Mintek
Time Mining and Processing (Pty) Ltd
Modular Mining Systems Africa (Pty) Ltd
Tomra Sorting Solutions Mining (Pty) Ltd
MSA Group (Pty) Ltd
TWP Projects (Pty) Ltd
Multotec (Pty) Ltd
Ukwazi Mining Solutions (Pty) Ltd
Murray and Roberts Cementation
Umgeni Water
Nalco Africa (Pty) Ltd
VBKOM Consulting Engineers
Namakwa Sands (Pty) Ltd
Webber Wentzel
New Concept Mining (Pty) Limited
Weir Minerals Africa (Pty) Ltd
Northam Platinum Ltd - Zondereinde
Xstrata Coal South Africa (Pty) Ltd
Department of Water Affairs and Forestry Deutsche Securities (Pty) Ltd Digby Wells and Associates Downer EDI Mining DRA Mineral Projects (Pty) Ltd Duraset E+PC Engineering and Projects Company Ltd
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APRIL 2014
The Journal of The Southern African Institute of Mining and Metallurgy
University of Pretoria Department of Mining Engineering Educating and leading mining engineers into the future Our broad-based curriculum emphasises the international engineering education model of conceptualising, designing, implementing and operating mines. The undergraduate programme is accredited by the Engineering Council of South Africa (ECSA).
to closure of mines, whether it be hard rock or coal deposits, surface or underground operations. The development of management/people skills is incorporated in some of the final year modules in the undergraduate programme.
We offer a range of programmes: • Undergraduate • Postgraduate • Research • Continued Professional Development (CPD) courses
Our portfolio of undergraduate, postgraduate and research degree programmes provide an excellent basis for a rewarding career in mining.
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A degree at the University of Pretoria covers all aspects of mining engineering from pre-feasibility assessments
For more information contact the head of department, Prof Ronny Webber-Youngman on 012 420 3763 or e-mail daleen.gudmanz@up.ac.za or visit our website www.up.ac.za.
Universiteit van Pretoria • University of Pretoria • Yunibesithi ya Pretoria Privaatsak • Private Bag • Mokotla wa Poso X20 Hatfield 0028 Suid-Afrika • South Africa • Afrika Borwa Tel: +27 (0) 12 420 4111 • Fax • Fekse: +27 (0) 12 420 4555
www.up.ac.za
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2014/01/29 12:35 PM