Saimm 201405 may

VOLUME 114

NO. 5

MAY 2014

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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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ii

MAY 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 R.J. Stewart

D. Tudor

Typeset and Published by 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

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.

NO. 5

MAY 2014

Contents Journal Comment

by P. Smith . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

v

President’s Corner

by M. Dworzanowski . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

vii

Special Articles South African Mineral Resource Committee (SAMREC): Re-write of the SAMREC Code (2014)

by K. Lomberg. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

viii

Base Metals papers Sulphuric acid plant water saving options

by R.J. Forzatti, I. Natha, L. Roux, and D.A. van den Berg . . . . . . . . . . . . . . . . . . . . . . . . . . . .

355

Challenges and successes at the Nkomati Nickel JV: pit-to product process improvements

by G. Cockburn . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

365

Evaluation of a Versatic 10 acid/Nicksyn™ synergistic system for the recovery of nickel and cobalt from a typical lateritic leach liquor

by A.C. du Preez and M.H. Kotze . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

375

General papers Evaluation of different adsorbents for copper removal from cobalt electrolyte

by V. Yahorava, M. Kotze, and D. Auerswald . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

383

Thermodynamic analysis and experimental study of manganese ore alloy and dephosphorization in converter steelmaking

by G. Chen and S. He. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

391

Atmospheric oxidative and non-oxidative leaching of Ni-Cu matte by acidified ferric chloride solution

by L.M. Sekhukhune, F. Ntuli, and E. Muzenda . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

401

Universities and decision-making: programme and qualification mix – four learning pathways

by W.P. Nel . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

411

Erratum: ‘A study on the effect of coke particle size on the thermal profile of the sinters produced in Esfahan Steel Company (ESCO)’,

by A. Dabbagh, A. Heidary Moghadam, S. Naderi, and M. Hamdi . . . . . . . . . . . . . . . . . . . . . .

418

VOLUME 114

NO. 5

MAY 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

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Editorial Consultant

VOLUME 114

Journal Comment

The Journal of The Southern African Institute of Mining and Metallurgy

probability that nickel lateritic ores will be increasingly exploited in the future in comparison to sulphide ores.

Although HPAL was first implemented many decades ago at Moa Bay in Cuba, the technology has faced many techno-economic challenges. One has been the difficulties associated with calcium as encountered in the Australian ‘dry lateritic’ projects during the 1990s.

This paper presents the results of laboratory test work which demonstrates that calcium co-extraction can be avoided in the extraction and stripping stages of the solvent extraction plant by judicious selection of reagents and operating conditions.

Evaluation of different absorbents for copper removal from cobalt electrolyte by V. Yahorava, et al.

Several fibrous ion exchangers were investigated and compared to the more conventional granular ion exchangers for the removal of copper from cobalt electrolytes. A comparison of the design parameters and indicative costs for the impurity removal process is presented for the two alternatives.

P. Smith

MAY 2014

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The 7th Southern African Base Metals Conference was held in Mpumalanga from 2 to 3 September 2013, with a visit to Nkomati Nickel Mine on 3 September. The Conference attracted 22 papers from sub-Saharan Africa (DRC, Zambia, Namibia, and South Africa) as well as from Finland and Australia. The scope of the papers was wide-ranging, including geology, engineering design, and process metallurgy. This edition of the SAIMM Journal includes a selection of four papers from the conference. It is regrettable that four papers were withdrawn from publication for a variety of reasons at the request of the authors.

Sulphuric acid plant water savings options by R.J. Forzatti, et al. considers strategies for saving water in sulphur-burning acid plants, although the principles also apply to acid plants burning sulphide minerals such as pyrite, sphalerite, etc. Acid plants basically involve exothermic reactions on a huge scale, and require extensive cooling. The paper presents an economic evaluation of various options to achieve this goal against various backgrounds of localized power and water costs. Given the current sensitivity to environmental factors throughout the world, this paper provides an important contribution to the debate. In general, mines tend to be found in remote locations where both power and water are at a premium and indeed, sometimes not available at all. The northern regions of Chile provide a good example of this with the challenges of the Atacama Desert.

Challenges and successes at the Nkomati Nickel JV: pit to product process improvements by G. Cockburn is of particular interest in that the deposit was first subjected to a feasibility study in the 1970s (INCO/Anglo American). Several subsequent attempts all failed to build a viable case. However, today there exists a successful operating mine with a growth profile that can only be described as spectacular – 10 000 t/month to 700 000 t/month. A mine-to-mill optimization programme is described along with the consequent benefits – significantly the effect of blasthole patterns and explosive powder factors on primary crusher feed size distribution and the critical role that a stable plant throughput plays in milling and flotation performance.

Evaluation of a versatic 10 acid/Nicksyn™ synergistic system for the recovery of nickel and cobalt from a typical lateritic leach liquor by A.C. Du Preez and M.H. Coetzee is a valuable contribution to high-pressure acid leach (HPAL) nickel technology, given the

â–˛

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MAY 2014

The Journal of The Southern African Institute of Mining and Metallurgy

T

he South African mining industry has been mainly associated with gold, which is understandable given that it was the gold mines of the late nineteenth century that were really the beginning of the industry as we know it. Currently platinum mining is making headlines regarding long strikes, and coal mining and Eskom are also much in the news. Diamonds have always featured in the media, to an extent that varies with time. However, base metals have never received any prominence, although they have formed part of the industry’s contribution for many decades. Copper, nickel, lead, and zinc are produced in South Africa.

Palabora Mining Company produces copper, Nkomati Nickel produces nickel, and Black Mountain produces copper, zinc, and lead. In addition, copper, nickel, and cobalt are by-products from the base metals refineries associated with the four major platinum producers.

Copper is produced in the form of cathodes from electrorefining and electrowinning. Nickel is produced in the form of metal cathodes by electrowinning, or as metal briquettes or nickel sulphate.

Cobalt is produced in the form of metal briquettes or cobalt sulphate. There is no primary production of lead metal, and there is no longer any primary production of zinc metal since the closure of Zincor at the end of 2011. With the exception of lead and zinc, the beneficiation of base metal ores in South Africa is thus well developed.

The beneficiation of base metal ores will always provide a significant challenge to extractive metallurgists – I speak from personal experience with many base metal projects in southern Africa. The flow sheet options are considerable, and in many instances mineral processing, pyrometallurgy, and hydrometallurgy need to be applied.

If we broaden our view of base metals to the southern African region, then we see a copper and cobalt industry of global significance. The Central African Copperbelt that spans Zambia and the Democratic Republic of Congo (DRC) is a world-class mineral province. Although mined tonnage does not compare with the copper porphyries of North and South America, the higher copper grades mean that actual copper production is not that far behind. Zambia and the DRC are 5th in global copper production when their output is combined. Beneficiation of copper ores, both sulphide and oxide, in Zambia and the DRC is well developed, with most of the copper being produced as cathode metal via electrorefining or electrowinning. There is a significant diversity of copper concentrators, smelters, and refineries within Zambia and the DRC. When the Nchanga tailings leach complex in Zambia was originally built close to 40 years ago it boasted the world’s largest copper solvent extraction and electrowinning plant. At Ndola in Zambia is one of the world’s few refineries processing copper refinery anode slimes, producing selenium, tellurium, silver, and gold as by-products.

The Copperbelt constitutes the world’s largest deposit of cobalt, which is associated with the copper in oxide and sulphide minerals. There are a number of cobalt plants which beneficiate oxide and sulphide cobalt concentrates into cathode metal via electrowinning. Zambia and the DRC produce about half the world’s cobalt.

There are nickel mines in Botswana and Zimbabwe, and zinc, lead, and copper mines in Namibia.

Zinc and lead are also produced on a small scale in Zambia and the DRC. This all highlights the extent of base metals production in southern Africa and illustrates why the SAIMM organizes a base metals conference every two years. The next conference will be held in Zambia in 2015, with a copper / cobalt theme. The event will be convened in conjunction with our Zambia and DRC branches, and will serve as a vehicle for promoting them, as well as provide a means to motivate the hosting of the International Copper Conference in Southern Africa.

entʼs d i s e Pr er Corn

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M. Dworzanowski President, SAIMM

South African Mineral Resource Committee (SAMREC) Re-write of the SAMREC Code (2014)

T

he mining industry is a vital contributor to national and global economies; never more so than at present with soaring demand for the commodities that it produces. It is a truly international business that depends on the trust and confidence of investors and other stakeholders for its financial and operational well-being. Unlike many other industries, it is based on depleting mineral assets, the knowledge of which is imperfect prior to the commencement of extraction. It is therefore essential that the industry communicates the risks associated with investment effectively and transparently in order to earn the level of trust necessary to underpin its activities. (CRIRSCO website) The SAMREC Code, which sets out minimum standards, recommendations, and guidelines for Public Reporting of Exploration Results, Mineral Resources and Mineral Reserves in South Africa, is being reviewed and improved to ensure that it remains relevant to the minerals industry and keeps abreast with recent developments. This revision is considered necessary because as the guidelines of the Code are used, various issues and practical realities have become apparent that require further guidance from the Code. This rewrite is designed to improve the Code and eliminate possible contradictory reporting practices, and align SAMREC with recent changes to international codes in keeping with international best practice.

The SAMREC Code is one of seven codes that are affiliated under the CRIRSCO family of reporting codes. As a result of the CRIRSCO/CMMI initiative, considerable progress has been made towards widespread adoption of globally consistent reporting standards. These are embodied in similar Codes, guidelines, and standards published and adopted by the relevant professional bodies around the world. The definitions in this edition of the SAMREC Code are either identical to, or not materially different from, existing international definitions. In recent years the Russian Code (NAEN) (2011) was added to the original Codes.

Various Codes have been revised and reissued – CIM of Canada (2010), PERC representing Europe (2013), JORC representing Australia and New Zealand (2012), and SME representing the USA (under review for issuing in 2014).

Various aspects of the Code remain unchanged. Because SAMREC is part of the CRIRSCO family, there are 15 core definitions e.g. Mineral Resource, Mineral Reserve etc. that are common between the international codes. These are not being changed. Rather, the guidance and interpretation is being improved so that the Code is relevant. The Code remains a guideline for minimum public reporting of Exploration Results, Mineral Resources, and Mineral Reserves. The desire of the SAMREC Working Group is that the Code is used for all forms of reporting of Exploration Results, Mineral Resources, and Mineral Reserves, both public and private. The principles that underpin the code remain Transparency, Materiality, and Competence.

The Code requires that anyone who uses the Code and asserts themselves as a Competent Person (CP) in accordance with the Code needs to have five years’ relevant experience and be registered with SACNASP or ECSA or be a member of GSSA, SAIMM, or PLATO or a recognized professional organization (RPO).

A body whose members put themselves forward as CPs is required to have a code of ethics and a disciplinary code.

Scientists working in South Africa are required to comply with the Natural Scientific Professions Act of 2003. However, where the SAMREC Code is used as the basis for a mineral resource or reserve declaration that falls outside of the jurisdiction of South Africa laws and the CP declares his/her membership of GSSA or SAIMM in support of the declaration, then these organizations require the CP to follow the newly instituted procedure.

Because the GSSA and SAIMM are not statutory bodies and represent broader interests than just minerals reporting, the GSSA and SAIMM have introduced by-laws that require individuals who utilize their membership as a credential for reporting purposes to notify the societies and subject themselves to a peer review prior to the publication of the work. This peer review entails confirming that they are members of the societies in the category they claim and have the necessary qualification and experience to undertake this assignment as a CP. However, this does not militate against the individual producing work that is substandard. Should the individual complete substandard work and a complaint is laid, they will be subject to the disciplinary process.

Issues regarding the rewrite are discussed at a monthly meeting of the SAMREC Working Group (WG) chaired by Ken Lomberg (ken.lomberg@coffey.com) and held on the last Thursday of each month at the Military Museum in Saxonwold. All interested parties are invited to participate. These meetings also provide an opportunity for industry to highlight aspects that may need to be reviewed or improved. We would like to encourage all interested parties to submit any issues relevant to the rewrite of the SAMREC Code via the SAIMM (sam@saimm.co.za) by 30 June 2014. The intention is to complete a draft for public comment by the end of Q3 2014.

Once a draft has been finalized it will be issued for comment prior to being ratified by the SAMREC/SAMVAL Committee (SSC). It is also the intention of the SAMREC WG to prepare a companion volume that would include the practical application of the Code and assist in providing a benchmark for all industry practices. This volume is likely to be produced after the launch of the Code as the proceedings of a SAMREC conference.

K. Lomberg

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MAY 2014

The Journal of The Southern African Institute of Mining and Metallurgy

PAPERS IN THIS EDITION These papers have been refereed and edited according to internationally accepted standards and are accredited for rating purposes by the South African Department of Higher Education and Training

Base Metals papers Sulphuric acid plant water saving options

by R.J. Forzatti, I. Natha, L. Roux, and D.A. van den Berg . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 355 This paper considers several alternative cooling water systems for a conceptual 2000 t/d sulphuric acid plant.

A proprietary design tool is used to compare design options on both an economic and a weighted sustainability scale.

Challenges and successes at the Nkomati Nickel JV: pit-to product process improvements

by G. Cockburn . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 365 The metallurgical, operational, and management challenges involved in a number of milling and flotation optimization initiatives at the Nkomati Nickel JV operation are discussed, together with the outcomes obtained.

Evaluation of a Versatic 10 acid/Nicksyn™ synergistic system for the recovery of nickel and cobalt from a typical lateritic leach liquor

by A.C. du Preez and M.H. Kotze. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 375 A synergistic solvent extraction system is evaluated on a laboratory scale for the recovery of nickel and cobalt from a synthetic lateritic sulphate leach liquor, without the co-extraction of calcium.

Evaluation of different adsorbents for copper removal from cobalt electrolyte

by V. Yahorava, M. Kotze, and D. Auerswald . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 383 Granular and fibrous ion exchangers are compared for the removal of copper from cobalt advance electrolyte. The results are used to size a full-scale operation and carry out a techno-economic evaluation of the two ion exchange systems.

General papers Thermodynamic analysis and experimental study of manganese ore alloy and dephosphorization in converter steelmaking

by G. Chen and S. He . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 391 In this study, the effects of slag compositions, slag amount, temperature, and the carbon content of steel on the manganese and phosphorus distribution ratios during converter steelmaking are analysed. The results of the research could be useful in deciding on the application of manganese ore in alloying and identifying the slagging regime in converter steelmaking.

Atmospheric oxidative and non-oxidative leaching of Ni-Cu matte by acidified ferric chloride solution

by L.M. Sekhukhune, F. Ntuli, and E. Muzenda . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 401 The atmospheric leaching of nickel from a copper-bearing matte by acidic ferric chloride solution was studied at the laboratory scale. Leaching was found to be diffusion-controlled, and took place via three separate mechanisms that occurred simultaneously throughout the process. Oxidative leaching yielded higher nickel recoveries than non-oxidative leaching.

Universities and decision-making: programme and qualification mix – four learning pathways

by W.P. Nel . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 411 Many factors have to be considered when deciding on a Programme Qualification Mix (PQM). This paper focuses on the Higher Education Qualifications Sub-Framework (HEQSF) requirements, Engineering Council of South Africa (ECSA) standards and registration, and how these, together with the various qualifications and educational Learning Programmes (LPs) provided for by the HEQSF, may impact on the PQM decision taken by engineering departments and schools at South African universities.

These papers will be available on the SAIMM website

http://www.saimm.co.za

Sulphuric acid plant water saving options

by R.J. Forzatti*, I. Natha†, L. Roux†, and D.A. van den Berg*

➤ Heat recovery acid coolers, which preheat boiler feed water.

Synopsis

Keywords sulphuric-burning acid plant, steam, power generation, cooling systems, sustainable design.

Introduction Conventional sulphuric acid plants require water for their cooling systems. Cooling is required to reject surplus heat not recovered as steam. Reducing the water consumption lowers the cost of sourcing reliable supplies of clean water as well as the cost associated with treating effluent streams. It also helps improve the sustainability of the acid plant operation by reducing the impact on surrounding communities.

This paper focuses on two broad categories of water-saving options: ➤ Pretreatment of the make-up water required for evaporative cooling systems ➤ Replacement of evaporative cooling with dry cooling.

The sustainability aspect of these options is analysed using Hatch’s 4 Quadrant design tool.

Technologies integrated with the acid plant design that recover heat from acid will reduce the overall cooling water demand (and hence make-up water consumption). For the purposes of this evaluation, the following acid plant technology options are not considered: ➤ HRS (by MECS) and HEROS (by Outotec), which produce useful lowpressure steam The Journal of The Southern African Institute of Mining and Metallurgy

Overview A conventional 2000 t/d sulphur-burning sulphuric acid plant is considered in this paper.

Process description Solid sulphur is delivered in bulk bags or containers and is stored and transferred to a melting and filtration circuit. The sulphurmelting system uses low- and mediumpressure (LP and MP) steam, usually provided from the acid plant steam system. Dirty molten sulphur is filtered to remove ash and other solid impurities. Molten sulphur is also sometimes received instead of solid sulphur if the sulphur source is nearby.

Clean sulphur is transferred to the acid plant where sulphuric acid is produced. The sulphur is burned in a furnace at approximately 1200°C in contact with dry air to produce SO2 gas (approx. 12 vol.%) (King, Davenport, and Moats, 2013). The SO2 gas is oxidized to SO3 gas in contact with a vanadium pentoxide-type catalyst. The SO3 is then absorbed and reacts with the aqueous component of strong sulphuric acid to produce H2SO4. Circulating and product acid cooling is achieved in heat exchangers supplied with cooling water, typically provided from an evaporative cooling tower.

The acid plant steam system is designed to recover the heat generated by the exothermic

* Hatch Associates, Perth, Western Australia. † Hatch Goba, Woodmead, South Africa. © The Southern African Institute of Mining and Metallurgy, 2014. ISSN 2225-6253. This paper was first presented at the, Base Metals Conference 2013, 2–4 September 2013, Ingwenyama Conference & Sports Resort, Mpumalanga.

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The production of sulphuric acid from sulphur generates heat. The majority of this heat is recovered as steam and is often used to generate electricity.

Heat not recovered as steam is rejected to cooling water systems. The design of the turbine, condenser, and cooling water systems impacts the overall water, energy, and environmental footprint of the plant. This review considers a conceptual 2000 t/d sulphuric acid plant with several alternative cooling water systems. The review utilizes Hatch’s 4 Quadrant sustainable design tool to compare the alternatives on both an economic and a weighted sustainability scale.

Sulphuric acid plant water saving options reactions within the acid plant. Heat is recovered from the sulphur burner off-gas via the production of saturated highpressure (40 barg or 60 barg) steam in a waste heat boiler.

Saturated steam from the boiler flows through a superheater to produce superheated steam, which is fed to a steam turbine generator to produce electricity.

Exhaust steam from the acid plant turbine is condensed, re-pressurized, and returned to the acid plant as boiler feed water. The turbine steam condenser (‘surface’ condenser) uses cooling water from an evaporative cooling tower. The evaporative cooling tower loses water through evaporation, drift (entrainment), and blowdown. A continuous supply of fresh water is required to make up for these losses.

Demineralized water is used as make-up for losses within the steam circuit (e.g. boiler blowdown and deaerator vent) and for dilution water within the acid plant.

Typical operating parameters for a 2000 t/d sulphurburning acid plant are shown in Table I.

Electricity generation Steam produced by the acid plant can be: ➤ Used to generate electricity in a steam turbine generator ➤ Supplied to other plant consumers (e.g. heating for hydrometallurgical equipment) ➤ Exported to other customers ➤ Condensed.

Sulphur-burning acid plants produce more electricity than they consume when all of the steam is sent to a steam turbine generator. This excess electricity can be: ➤ Used to operate other facilities within the plant ➤ Sold to the market.

Water balance The water balance for a conventional 2 000 t/d sulphurburning acid plant based on evaporative cooling is given in Table II. The following is noted: ➤ The single largest water loss is due to water evaporation in the cooling tower ➤ Other losses include drift and blowdown. The blowdown indicated is calculated based on three cycles of concentration, assuming fresh water has a total dissolved solids (TDS) of 300–400 ppm.

Water-saving options Make-up water pretreatment Treatment of make-up water to the cooling tower can be used to change the water chemistry to achieve higher cycles of concentration, thereby reducing blowdown.

Softening Softening of the cooling water make-up can be used to remove several dissolved salts that cause scale formation such as calcium, magnesium, barium, strontium, and iron.

Other scale-forming components, such as silica, are not removed.

A water softener consists of a vessel filled with cationic resin that exchanges (removes) the dissolved species from the water and replaces these with sodium. Cooling systems fed with high hardness water sources will benefit most from having the make-up water softened.

As an example, a water source with a feed hardness of approximately 300 mg/L (expressed as CaCO3), pH of approximately 8, and alkalinity of approximately 300 mg/L as CaCO3 might normally be concentrated three times; a cycles of concentration (COC) of 3. The scaling tendency of this water, at a COC of 3, is within the typical range that can be managed with a scale inhibitor. This same make-up water source could

Table I

Typical 2 000 t/d acid plant operating parameters Parameter

Units

Value

Acid production (100% H2SO4 basis)

t/d

2 000

Sulphur consumption

t/d

660

Steam production (superheated)

t/h

110

Electricity generation (steam turbine)

MWe

23

Electricity consumption

MWe

5

Main acid coolers

106 kJ/h

144

Product acid cooler

106 kJ/h

7

Turbine surface condenser

106 kJ/h

234

Other coolers

106 kJ/h

11

Nominal cooling duty

106 kJ/h

396

Design cooling duty*

106 kJ/h

468

Cooling requirements

*Design cooling duty includes an additional 72 x 106 kJ/h installed capacity for when the steam turbine is bypassed

Table II

Sulphuric acid plant water balance Inputs

Inflow H2O (t/h)

Outputs

Outflow H2O (t/h)

Air moisture (to sulphur burner) Cooling tower make-up water Water to demin plant (for acid dilution)

4 220 24

Total Inputs

248

Steam deaerator vent Water in product acid Water converted to H2SO4 by reaction Cooling tower evaporation and drift loss Cooling tower blowdown Other effluent (steam system and demin plant) Total outputs

2 1 15 147 73 10 248

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Sulphuric acid plant water saving options be concentrated more than seven times if first softened, representing a reduction in cooling water make-up of approximately 20%.

Filtration Filtration to remove suspended solids from water can be applied to either the entire fresh water make-up stream for the acid plant or to the cooling tower water recirculation stream. Filtration of the cooling circuit make-up water is generally considered where there is a high level of suspended solids in the feed. These solids, if not removed, can cause fouling within the cooling water circuit, which lowers cooling efficiency and increases the pressure drop through the piping system. The solids can also accelerate corrosion within the water circuit if they are abrasive. Unfiltered particles can serve as nucleation sites for biological growth. Filtration is required ahead of a cooling water softening system.

It should be noted that the majority of suspended solids in the cooling circuit are generated within the cooling circuit rather than introduced in the make-up water, as the cooling water is in contact with surrounding air in open-circuit evaporative cooling towers. Internal sources of solids include pipe corrosion products, biological growth material, and dust introduced from the air as it contacts the water in the cooling water tower. For this reason there is often more merit in filtering the cooling water itself rather than the make-up water alone.

Demineralized water treatment Demineralized water is used as make-up for losses within the steam circuit (e.g. boiler blowdown and de-aerator vent) and for dilution water within the acid plant. Typical demineralized water system configurations include: ➤ Reverse osmosis (RO) only ➤ Ion exchange (IX) with a decarbonator tower ➤ RO followed by polishing IX. Waste generation as a percentage of feed is typically 30%.

The selected configuration is dependent on the site raw water quality, and can be optimized to provide water savings. These savings will, however, be small compared to potential savings in the cooling system.

Cooling technologies The following cooling technologies are discussed: ➤ Evaporative cooling towers ➤ Dry cooling technologies ➤ Hybrid cooling towers.

➤ Better operating procedures and equipment to monitor and control blowdown ➤ Use of high-efficiency drift eliminators and equipment to recapture drift ➤ Optimizing the selection and amount of fill inside the tower, which affects the heat transfer efficiency of the tower ➤ Automatic blowdown based on conductivity to avoid unnecessary blowdown in cases where the feed water quality is better than initially anticipated ➤ Minimizing unintentional water losses from leaks or overflow (i.e. faulty level control resulting in addition of excess make-up water) ➤ Special tower design considerations to reduce particulates, debris, and cooling water exposure to sunlight.

These advances in cooling tower design and control have resulted in minor water savings. Fundamentally, the cooling is provided through the evaporation of water, and hence there is an inherent loss of water when adopting this technology.

Dry cooling technologies Dry cooling technologies work by heat exchange to air and do not rely on the evaporation of water to provide cooling.

Applicable technologies for a sulphur-burning acid plant include: ➤ Air-cooled condenser (ACC) on the steam turbine exhaust, ➤ Fin-fan coolers to supply cooling water to the absorbing acid coolers.

Dry cooling technologies are dependent on the difference between ambient temperature and the cooling water temperature. In locations with high ambient temperatures, the temperature difference will be lower, leading to significantly increased dry cooling unit size.

Air-cooled condenser (ACC) An ACC (Figure 1) is comprised of finned tube bundles grouped together into modules and mounted in an A-frame configuration on a concrete or steel support structure. Steam from the turbine exhaust enters the top of the condenser via a steam duct and manifold. Steam flows downward through two or three rows of finned tubes. Condensate is recovered inside bonnet header boxers connected to a hot water tank.

The axial-flow forced-draught fan is fixed in the module and forces the atmospheric cooling air across the condensate area of the fin tubes.

Evaporative cooling towers Evaporative cooling tower designers have identified many ways to reduce the overall water losses from these systems.

Some of these include (EnduroSolv, 2012):

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Figure 1—Schematic of an ACC (SPX, 2012) VOLUME 114

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➤ Optimizing water chemistry to reduce scaling, corrosion, and biological growth, subsequently increasing the cycles of concentration and decreasing blowdown. This includes the use of automated chemical dosing systems

Sulphuric acid plant water saving options Dry cooling in an ACC requires a significant temperature difference to provide adequate heat exchange to the surrounding air. Typically, the cooling water supply temperature will be 25°C to 30°C higher than the ambient air temperature. This results in a higher condenser outlet temperature which in turn raises the condenser pressure, causing the steam turbine to operate less efficiently. An ACC can also be impacted by wind direction and speed as well as proximity to large buildings. More recent advancements in ACC technology include (Mortensen, 2011; Maulbetsch, DiFilippo, and Zammit, 2008): ➤ Wind guide vane technologies to mitigate wind impacts including walls, screens, lips, and louvers. CFD wind flow modelling is also used to optimize the location and arrangement of the ACCs ➤ Improved finned tube bundle designs for higher heat transfer efficiency and lower pressure drop ➤ Pre-cooled ACC, which uses the evaporative cooling effect of a fog spray into the upstream side of the ACC fans. The expected water consumption is approximately 75% less than equivalent evaporative-only cooling. The advantage of this system is that it reduces air temperatures to the fans on very hot days.

The application of an ACC for cooling of turbine exhaust is widely adopted on many steam turbine systems. Eskom, the South African power utility, has adopted the largest ACCs currently in operation in the world for the Matimba, Kendal, and Majuba power stations (Eskom, 2010).

for the drying and absorption sections of an acid plant because the absorbing acid heat exchangers target approximately 70°C. The product acid heat exchangers target 35–40°C, which cannot be consistently achieved in most locations using fin-fan coolers.

Hybrid cooling towers Hybrid cooling towers have an air-to-air (dry cooling) section and evaporative cooling section operating in series. As shown in Figure 3, heated cooling water first passes through the dry section, where part of the heat load is removed by an air current, typically induced via fans. After passing the dry section, water is further cooled in the wet section of the tower, which can be cooled in a conventional open evaporative circuit or closed circuit (tubes are cooled with water on the outside).

The resulting heat transfer performance is similar to a wet cooling tower, with the dry cooler providing the advantage of protecting the working fluid from environmental exposure and contamination. Depending on the hybrid tower configuration, the water consumption lies between the wet and dry circuit options reported in this evaluation.

Cooling technology and steam turbine electricity generation The turbine exhaust cooling system performance directly affects the amount of electricity produced by the steam turbine generator. The lower the condenser outlet

Fin-fan coolers Fin-fan coolers (as depicted in Figure 2) include one or more bundles of finned tubes connected by headers with an airmoving device such as an axial fan located above (induced draught) or below (forced draught) the tube bundle. Cooling water flows through the tubes and heat is exchanged to ambient air. The fin-fan circuit uses demineralized quality water and is closed-loop (not open to atmosphere), eliminating the need for a continuous water supply.

Fin-fan coolers require a significant temperature difference to provide adequate heat exchange to the surrounding air. Typically, the cooling water supply temperature will be 15°C higher than the ambient air temperature. A fin-fan cooler can be used to provide cooling

Figure 2—Fin-fan cooler (Wilson, 2011)

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Figure 3—Hybrid cooler examples (after EPRI, 2002) The Journal of The Southern African Institute of Mining and Metallurgy

Sulphuric acid plant water saving options temperature, the lower the condenser outlet pressure and turbine exhaust back-pressure. A lower turbine exhaust back-pressure increases turbine output.

The impact of increasing ambient temperature (dry-bulb temperature) on the turbine output is shown in Figure 4, similar to the loss of electricity generation reported by others (US Department of Energy. 2009). It compares the base case (turbine surface condenser on evaporative cooling) to the turbine ACC option. For both options, the equipment is sized to remove the full heat load over the full ambient temperature range. As can be seen, with an ACC, the power generation is lower because an ACC runs at a higher temperature than a surface condenser.

Sustainability impacts Sulphur-burning acid plant emissions include: ➤ Gaseous emissions—sulphur dioxide, nitrogen oxides, and acid mist in tail gas ➤ Liquid effluents—waste heat boiler and cooling circuit blowdown, demineralized water treatment plant waste, plant washings, spillages and leakages ➤ Solid effluent—sulphur filter cake residues and spent converter catalyst ➤ Noise pollution—main blower and turbine.

Summary The following cooling circuit water-saving options are compared in this paper: ➤ Base case – evaporative cooling (no pretreatment) ➤ Evaporative cooling (with pretreatment) ➤ Dry cooling (no pretreatment) with the following variants: – ACC on turbine exhaust – ACC on turbine exhaust, and fin-fans on absorbing acid circuit.

Figure 4—Impact of ambient temperature on turbine electricity generation

These options are shown in the schematic in Figure 5.

All options assume steam is generated in the sulphur burner waste heat boiler and electricity generated in a steam turbine generator. Where dry cooling options are considered, the balance of cooling is provided by evaporative cooling.

Table III summarizes the make-up water consumption and heat removal duties of the cooling circuit options.

Evaluation of options The Hatch 4QA approach compares the economic and sustainability impacts of alternative project options.

Economic impact Order–of-magnitude capital and operating cost estimates were developed for each cooling system option and compared to the base case (evaporative cooling with no pretreatment). A summary of the comparison is presented in Table IV.

Figure 5—Water saving options schematic

Table III

Summary of cooling circuit options Parameters

Units

Evaporative cooling

Dry cooling (no pretreatment)

No pretreatment

Pretreatment included

ACC

ACC +Fin-fan

Make-up water consumption Make-up water

t/h

220

172

86

7

Heat removal duty (nominal) Evaporative cooling tower ACC Fin-fan Total (nominal) Total (design)*

106 kJ/h 106 kJ/h 106 kJ/h 106 kJ/h 106 kJ/h

396 0 0 396 468

396 0 0 396 468

162 234 0 396 468

18 234 144 396 468

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*Design cooling duty includes an additional 72 x 106 kJ/h capacity for when the steam turbine is bypassed

Sulphuric acid plant water saving options Table IV

Cooling system capital and operating cost comparison Cost parameters

ACC

ACC + Fin-fan

No pretreatment

Base case With pretreatment

No pretreatment

No pretreatment

1.00 1.00

1.21 1.18

1.33 0.99

1.43 0.87

CAPEX (relative to base case) OPEX (relative to base case)

Table V

Sustainability criteria values for cooling system options Sustainability criteria values for

Base case

cooling system options Water intensity Power intensity Footprint intensity Waste intensity

m3/t acid kWh/t acid m2/t acid m3/t acid

ACC

ACC + Fin-fan

No pretreatment

With pretreatment

No pretreatment

No pretreatment

2.6 36 10 0.9

2.3 38 12 0.5

1.0 44 18 0.3

0.1 48 23 0.03

Commitments by major corporations as well as government regulatory requirements have resulted in the development of several new and cost-effective technologies to efficiently reduce gaseous emissions from an acid plant.

Noise pollution has been addressed with suitable sound reduction measures such as acoustic insulation, enclosures, and silencers.

The treatment and disposal of liquid and solid effluents is facing more stringent controls through public awareness and government regulations. Reduction in water consumption is potentially the greatest beneficial impact on local communities and the environment. Furthermore, a reduction in the cooling water blowdown will reduce the plant effluent and ultimately reduce the impact on the overall plant effluent catchment area.

Water, footprint, power, and waste Four sustainability criteria were identified to quantitatively compare the different water saving options, namely: water intensity, power intensity, waste intensity, and footprint. A comparison of these criteria is given in Table V. Water and waste intensity are calculated from the mass balance; power and footprint values are estimated from recent project experience. The values shown are for the cooling system only and exclude the criteria associated with the remainder of the acid plant.

The following is noted with respect to each of the intensity factors: ➤ Water intensity – Pretreatment provides water savings when applied to the base case – Dry cooling options provide the lowest overall water consumption. ➤ Power intensity – Evaporative cooling includes power to operate the cooling tower fans and the cooling water supply pumps

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– Base case with pretreatment has a slightly higher power usage due to more pumping required between upfront unit operations – Dry cooling includes power to operate the fans only. Although electricity consumption is lower for dry cooling options, the power intensity shown in Table V is higher because it has been calculated taking into account a reduction in turbine electricity generation of 1.5 MWe (see previously). ➤ Footprint intensity – The footprint of the base case with pretreatment is comparable with the base case – Dry cooling options require more footprint for the same cooling duty. ➤ Waste intensity – Evaporative cooling options generate more blowdown and therefore increased waste for effluent treatment – There is a large reduction in waste generated when pretreatment is included ahead of the evaporative cooling tower.

Four Quadrant Analysis Hatch developed the Four Quadrant Analysis (4QA) approach to compare project options using the economic and sustainability impacts. The 4QA tool plots each option compared to a base case: ➤ The x-axis is a cost ratio, with a lower cost ratio representing a lower cost option relative to the base case ➤ The y-axis is a sustainability ratio, which compares the intensity of the option to the base case. A lower sustainability ratio is preferred, which indicates a lower impact on the environment.

The cost ratio (CR) is calculated for each option, relative to the base case (BC), using the following equation: The Journal of The Southern African Institute of Mining and Metallurgy

Sulphuric acid plant water saving options ➤ Low water cost (US$0.2 per m3)—typical for locally available water source of good quality (e.g. dam located close to plant) with no additional extraction charges ➤ Average water cost (US$1.0 per m3)—typical for water sources located a reasonable distance from the plant, requiring minor infrastructure to be built and some minor water treatment on site (e.g. sand filtration) ➤ High water cost (US$3.0 per m3)—typical for water sources located at a considerable distance or water of poor quality requiring significant treatment (e.g. reverse osmosis). High water cost would also apply for water that is local and of good quality, but with a high extraction charge.

[1]

The CR is the sum of the annual operating cost and the annual capital cost repayments, based on a nominal 5-year repayment (compounded monthly). Any credits received for selling electricity to the market have not been included into the evaluation.

The sustainability ratio (SR) is calculated for each option, relative to the base case (BC), using the following equation:

[2]

Table VIII summarizes the cost ratios for the water cost sensitivity analyses. The sustainability ratios remain unchanged.

The sensitivity analysis shows that:

The weightings can be adapted based on the general importance of each criterion. The weightings used in this evaluation (Table VI) have placed a high importance on water intensity as many plants strive to reduce water consumption.

The relative weightings will be site-specific, for example arid locations may consider water impacts more important and footprint less important.

The relative cost ratio and sustainability ratio of the water- saving options are given in Table VII, and the 4QA plot is in Figure 6.

The 4QA shows that for typical site locations:

➤ Reduced water costs (US$0.2 per m3) increase the cost ratio of dry cooling options, making them unfavourable compared to the base case

Table VI

Weightings of sustainability criteria Sustainability criterion

Weighting

Water intensity Power intensity Footprint intensity Waste intensity

➤ Base case with pretreatment has a better sustainability ratio than the base case; however, the cost ratio will increase by 19%. This is mainly due to the high operating cost associated with reagent consumption for cationic resin regeneration. Alternative pretreatment options can be investigated with lower reagent usage and allowing increased cycles of concentration in the cooling tower ➤ Dry cooling offers considerable improvements to the sustainability ratio, but the cost ratio will increase by 10% for ACC and 4% for ACC and fin-fan.

40% 15% 20% 25%

Sensitivity analysis The cost and sustainability ratios can be affected by several factors, some of which are briefly considered in the following sensitivity analyses.

Water costs Fresh water supply costs are location-dependent. The relative water supply costs used for this sensitivity are based on:

Figure 6—Hatch 4Q Analyses (average fresh water and power cost)

Table VII

Ratio

Cost ratio Sustainability ratio

ACC

ACC + fin-fan

No pretreatment

Base case With pretreatment

No pretreatment

No pretreatment

1.00 1.00

1.19 0.87

1.10 0.77

1.04 0.67

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Sustainability and cost ratio summary

Sulphuric acid plant water saving options Table VIII

Cost ratios for water cost sensitivity Cost sensitivity

Base case

Low fresh water cost ratio Average fresh water cost ratio High fresh water cost ratio

ACC

ACC + fin-fan

No pretreatment

With pretreatment

No pretreatment

No pretreatment

1.00 1.00 1.00

1.24 1.19 1.10

1.22 1.10 0.91

1.22 1.04 0.77

ACC

ACC + fin-fan

No pretreatment

With pretreatment

No pretreatment

No pretreatment

1.00 1.00 1.00

1.21 1.19 1.13

1.08 1.10 1.13

1.00 1.04 1.14

US$0.2 per m3 US$1 per m3 US$3 per m3

Table IX

Cost ratios for electricity cost sensitivity Cost sensitivity

Base case

Low electricity cost ratio Average electricity cost ratio High electricity cost ratio

US$0.05 per kWh US$0.1 per kWh US$0.3 per kWh

➤ Increased water costs (US$3 per m3) reduce the cost ratio of dry cooling options to less than the base case (by as much as 23%) ➤ This supports the general observation that as water extraction costs increase, dry cooling options are preferred ➤ The base case with pretreatment has the highest overall cost ratio, which corroborates the data in Table VII. At reduced water costs (US$0.2 per m3) the base case with pretreatment compares well with the dry cooling options, but becomes the least favourable overall at increased water costs (US$3 per m3).

The 4QA was updated with the low and high water costs in Figure 7.

Electricity costs Electricity supply costs are also location-dependent. The relative electricity supply costs used for this sensitivity are based on:

the dry cooling options at increased electricity costs; however, it becomes the least favourable at the lower electricity costs.

The 4QA was updated with the low and high electricity costs in Figure 8.

Sustainability criteria weightings The sustainability criteria weightings can be adjusted to suit the plant location and requirement for generating electricity, thereby impacting the sustainability ratio. As an example, the weightings can be adjusted as water or power becomes more or less important to the local community. Furthermore, additional sustainability criteria can be included, such as: ➤ Specific reagent consumption (e.g. high RO membrane costs) ➤ Downstream impact and stewardship (qualitative) ➤ Operability and maintainability (qualitative)

➤ Low electricity cost (US$0.05 per kWh): for locations with abundant low cost electricity, e.g. hydroelectricity ➤ Average electricity cost (US$0.1 per kWh): for locations with a typical mixed electricity supply, e.g. a mix of coal, renewable, and gas-fired power stations ➤ High electricity cost (US$0.3 per kWh): for locations where electricity is generated on site, e.g. local diesel or gas-fired generators.

Table IX summarizes the cost ratios for the electricity cost sensitivity analyses.

The sensitivity analysis shows that: ➤ Lower electricity costs have a minor impact on the 4QA plot ➤ At increased electricity cost, the advantage of evaporative cooling is clear, due to the increased turbine electricity output ➤ The base case with pretreatment compares well with

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Figure 7—Hatch 4Q Analyses for water cost sensitivity The Journal of The Southern African Institute of Mining and Metallurgy

Sulphuric acid plant water saving options is generally lower for dry cooling, due to lower water consumption and similar power consumption to evaporative cooling ➤ Acid plant electricity generation capacity is lower with dry cooling options and is worsened during higher ambient temperature conditions. For large acid plants generating electricity that is sold to market, this will adversely impact plant revenue.

Acknowledgements The authors wish to acknowledge the contributions of Dr Matthew King and Rusi Kapadia, and Hatch for permission to publish this paper.

Figure 8—Hatch 4Q Analyses for electricity cost sensitivity

References ENDUROSOLV. 2012. Cooling Tower Water Saving Strategies, http://endurosolv.com/pdf/cooling_tower_savings_strategies.pdf [Accessed 25 Apr. 2014].

➤ Social, for example using local labour (qualitative) ➤ Government and externals (qualitative) ➤ Emissions (qualitative).

The 4QA approach is flexible and can be customized and adapted to meet specific project criteria.

EPRI. 2002. Comparison of Alternate Cooling Technologies for California Power Plants: Economic, Environmental and Other Trade-offs. Electric Power Research Institute, Aplo Alto, CA, and Californian Energy Commission, Sacramento, CA. ESKOM. 2010. Factsheet, General Communication CO 0005, Revision 7, October

The Hatch 4QA approach compares the economic and sustainability impacts of alternative project options. It can be used for technology and site selection from concept through to feasibility studies and beyond. The 4QA additionally serves as a risk management tool to quantify the impacts of varying sustainability criteria and input costs.

For the water-savings options considered in this paper, it is the acid plant location that largely determines the sustainability and cost ratios. Key findings include: ➤ The cost ratio of evaporative cooling is generally lower, provided there is good-quality and low- to mediumcost water available. The sustainability ratio is generally higher due to the high water consumption, which can make evaporative cooling unfavourable even at sites with low water costs (Maulbetsch, DiFilippo, and Zammit, 2008) ➤ The cost ratio of the base case with pretreatment is the highest overall due to the high reagent usage.

Optimizing the make-up water chemistry by adjusting the pH and adding scale inhibitors might be a more efficient way of increasing the cycles of concentration, but needs to be investigated on a case-by-case basis ➤ Reverse osmosis (RO) can also offer water savings, but these could be offset by the RO waste generation ➤ The sustainability ratio of the base case with pretreatment is lower than without pretreatment due to the decreased water usage ➤ Dry cooling options have a higher capital outlay, but can have lower operating costs in locations where water extraction costs are high. The sustainability ratio The Journal of The Southern African Institute of Mining and Metallurgy

2010. http://www.eskom.co.za/AboutElectricity/FactsFigures/Documents/CO000 5CoolingTechniquesRev10.pdf [Accessed 12 June 2013].

KING, M.J., DAVENPORT, W.G., and MOATS, M.S. 2013. Sulphuric Acid Manufacture: Analysis, Control and Optimization. 2nd edn. Elsevier, Burlington, MA. ISBN: 978-0-08-098220-5.

MAULBETSCH, J.S., DIFILIPPO, M.N., and ZAMMIT, K.D. 2008. spray cooling for performance enhancement of air-cooled condensers. EPRI Advanced Cooling Workshop, Carolina.

MORTENSEN, K. 2011. Improved Performance of an Air Cooled Condenser (ACC) Using SPX Wind Guide Technology at Coal-Based Thermoelectric Power Plants. US Department of Energy. http://www.netl.doe.gov/File%20Library/Research/Coal/ewr/water/Proj51 9.pdf [Accessed 24 April 2014].

SPX. 2012. Air Cooled Condensers. www.spx.com/en/assets/pdf/A4_ACC12.pdf [Accessed 18 June 2013].

US DEPARTMENT OF ENERGY. 2009. Concentrating Solar Power Commercial Application Study: Reducing Water Consumption of Concentrating Solar Power Electricity Generation. Report to Congress in response to Energy Independence and Security Act of 2007. (Pub. L. No. 110-140), http://www.nrel.gov/csp/pdfs/csp_water_study.pdf [Accessed 25 April 2014].

WILSON, B. 2011. Detail Engineering and Layout of Piping Systems. 1st edn. On Demand Books, New York. ISBN : 9781926633183.

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Ceramite® Extreme wear and

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Typical property ranges Density = 2.6 to 2.8 kg/dm3 Cold crushing strength = 70 to 220 MPa (20 to 1500ºC respectively) Flexural strength = 7 to 28 MPa (20 to 1500ºC respectively) Thermal conductivity range choice (0,6 to 14 W/mK)

thermal conductivity range Zero wear after 3 years

Minimal wear after 5 years

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Challenges and successes at the Nkomati Nickel JV: pit-to product process improvements

by G. Cockburn*

Nkomati Nickel JV exploits the ores of the Uitkomst Complex near Machadodorp in the Waterval Boven district in South Africa’s Mpumalanga Province.

Due to factors such as the remote location, stellar growth in production, opencast mining methods, and ore characteristics, a number of innovative processing options were selected.

Nkomati has undertaken numerous initiatives over the last few years to improve plant running times, metallurgical performance, and operational profitability. Great emphasis has been placed on effectiveness of management control systems. A number of initiatives such as short interval control and time–in-state metrics have been implemented. A focus on improvement on availability and asset utilization of key items of equipment has been particularly effective.

While the ores are remarkably similar to the Merensky and UG2 reefs, the relatively high base metal sulphide content and mineralogical characteristics make metallurgical treatment somewhat different to the ores of the Bushveld Complex. The low head grades and flotation kinetics distinguish Nkomati from other base metal operations. Numerous milling and flotation optimization initiatives have resulted in dramatic improvements in throughput, recoveries, and concentrate grades.

This paper discusses the metallurgical, operational, and management challenges and the outcomes obtained.

Keywords Uitkomst, Nkomati Nickel, sulphide mineralization, grade control, ore fragmentation, problem solving methodologies.

Introduction Nkomati Nickel JV has experienced a phenomenal growth rate over the last few years, from a 10 kt/month operation in 2006 to a 700 kt/month complex in 2013. This growth required the re-engineering of virtually every aspect of the operation, from mining new ore types with new methods, ore preparation and processing, to tailings deposition.

Many alternative processing methods were considered before the current flow sheets were adopted. Production of separate nickel and copper concentrates through successive selective flotation, Activox® leaching of concentrates, and local smelting of concentrates were assessed among other options.

Ultimately, economic factors resulted in the current circuit choices. This expansion occurred against the backdrop of the ongoing The Journal of The Southern African Institute of Mining and Metallurgy

The Uitkomst deposit Nkomati Nickel JV exploits the Uitkomst deposit in South Africa’s Mpumalanga province, in the mountains between Waterval Boven, Machadodorp, and Badplaas (Figure 1).

The orebody is an early-age (2044 Ma) Bushveld layered lenticular mafic-ultramafic intrusion into the basal sediments of the Transvaal Supergroup, approximately 9 km long and 1500 m wide (Figure 2). The deposit dips north-east at about 4 degrees. The deposit was exploited by AngloVaal with various partners since the early 1990s. Nkomati Nickel JV is a 50/50 partnership between African Rainbow Minerals and Norilsk Nickel Africa.

The orebody has multiple zones of sulphide mineralization: ➤ MSB: Massive Sulphide Body with Ni grades in excess of 2%. Mined since 1997, now mined out ➤ MMZ: Main Mineralized Zone. Head grades 0.3–0.7% Ni, approximately 0.37 % Ni average ➤ PCMZ: Peridotitic, Chromititic Mineralized Zone. Chrome-rich ore with grades of 0.2–1% Ni, 0.23% Ni average, chrome grades of 10-15% Cr2O3. ➤ Massive chromitite (often called PCR): stockpiled for a separate chrome washing plant, currently mothballed ➤ Basal Mineralized Zone: unexploited at present.

* Nkomati Nickel JV. © The Southern African Institute of Mining and Metallurgy, 2014. ISSN 2225-6253. This paper was first presented at the, Base Metals Conference 2013, 2–4 September 2013, Ingwenyama Conference & Sports Resort, Mpumalanga.

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global economic crisis, with persistent low metal prices.

This paper discusses many of the efforts that contributed the turnaround of Nkomati Nickel. Emphasis is placed on the challenges and successes at the MMZ plant, though the PCMZ plant showed similar improvements.

Challenges and successes at the Nkomati Nickel JV: pit to product process improvements

Figure 1—Location of Nkomati Nickel JV

Figure 2—Idealized cross-section of Uitkomst deposit

Figure 3—Life-of-mine ore production profile

Ore production Currently only the MMZ and PCMZ ores are mined. Ore production from the open pit is approximately 650 kt/month, of which approximately 300 kt is PCMZ and 350 kt MMZ. The MMZ is also mined in the underground mining section,

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producing approximately 50 kt per month by bord and pillar and longhole open stoping methods.

It must be noted that the production profile (Figure 3) is routinely optimized and updated as models are tuned and improved based on the outcomes of the RC drilling programme discussed below.

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Challenges and successes at the Nkomati Nickel JV: pit to product process improvements Mineralogy The pyroxenites and peridotites of the Lower Pyroxenite that hosts the MMZ consist mainly of clinopyroxene, olivine, and plagioclase. Hydrothermal action has resulted in extensive alteration of these minerals to amphibole, serpentine, biotite, and talc. Contamination of the ultramafic suite by country rocks accounts for the most of the calcite, dolomite, quartz, and plagioclase. Talc content is highly variable within the deposit and irregular (Brits, 2008) Highly altered talc-rich zones are often associated with pyrite-rich zones, and the high flotation kinetics of both these minerals complicates the flotation process, resulting in lower pentlandite recoveries, dilution of the concentrate with pyrite, and reduced concentrate quality due to higher MgO levels.

Nickel is mainly contained within pentlandite, although a significant proportion (as much as 15%) occurs in solid solution within pyrrhotite and 1–2% within chlorite. Copper occurs almost exclusively as chalcopyrite, with some occurring as bornite (1–2%).

The MMZ in many ways resembles Merensky Reef, although it contains substantially lower platinum group metals (PGMs) and higher base metal sulphides (typically 5–8%) with traces of PGMs (1 g/t head grade, predominantly Merenskyite). The PCMZ resembles the UG2 Reef, with chrome grades of 7–15% Cr2O3. From the geologist’s perspective, the ores are effectively the same with the exception of the chrome grades. The boundaries of the two ore types are not clearly delineated, making segregation of ore and prevention of cross-contamination challenging.

From a processing perspective, however, the ores are significantly different. The target liberation grind for MMZ ore is 67% -75 μm, although recoveries are relatively insensitive to grind. PCMZ ore is extremely sensitive to grind, with a target grind of 80% -75 μm, and drastic losses in recovery occur at lower grind values. Misplaced ore thus directly affects plant performance.

Grade control An extensive reverse circulation (RC) drilling programme at an initial hole spacing of 25 m × 25 m, and subsequently at 12 m × 12 m, has greatly enhanced the ability to model the orebody and so allow far better head grade control.

This is critical, considering the variability of grades within the orebody, and the fact that a substantial amount of PCMZ ore in particular is below economically viable grade.

Management of the resource is thus a vital aspect of maximizing the value of the mine. RC drilling data and the resource models derived from it are extensively used in mineto-mill reconciliations as well.

The Metso 54 × 75 Mk2 gyratory crusher was initially viewed as something of an Achilles’ heel of the operation.

The crusher suffered numerous breakdowns and trips and became the major process bottleneck. Although designed with an F80 of 450 mm and an F 100 of 1000 mm and a feed rate of 1600–1800 t/h, rocks substantially larger than design were routinely crushed, resulting in trips and mechanical failures. Tramp steel was also a major contributor to downtime.

Great focus was placed on preventing large rocks from entering the crusher. A ‘SPLIT’ camera and image analysis systems were introduced to monitor and record the size of rocks on trucks prior to tipping, with an additional system monitoring rock size during tipping. These systems provide a vital service in monitoring crusher feed PSD (Figure 4). Large rocks are prevented from entering the crusher largely through visual observation by control room operators, who reject truckloads with large rocks.

Interestingly, analysis of data indicated that highamperage trips (and damage) on the crusher were not caused by large rocks alone. Correlation of SPLIT rock size images, vibration, and ampere readings indicated that smaller football-sized rocks, mostly from re-broken boulders fed to the crusher in the absence of fines, were as much of a challenge as very large rocks. This was exacerbated by wear on the liners, where the crusher cavity would wear to a ‘hockey stick’ shape, trapping critical-sized rocks, akin to bearings in a race.

Key to improving throughput and availability was the implementation of strict planned maintenance systems. Key performance indicators such as overall equipment efficiency (OEE) were introduced and proved very effective in monitoring the actual crusher performance. OEE is calculated from actual tons processed divided by the theoretical maximum the equipment can process over the period of consideration. This measure cuts out all the clutter and confusion of allocation of downtime, and provides a ‘bottom line’ performance value.

The overland conveyor system capacity was increased to accommodate the increase in crushed tonnages from 1300 to 2000 t/h. This necessitated the installation of larger head pulleys, shallower troughing angles on belts, and faster belt speeds to reduce persistent belt splice failures. Attention to best practices in splicing the steel-cored belts was vital to increase availability.

Processing challenges Primary gyratory crusher

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Figure 4—PSD analysis of crusher feed using the SPLIT camera system VOLUME 114

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With the open pit supplying the vast bulk of the ore, a primary gyratory crusher at the pit was selected with overland conveyors to transport to the two plants. Loading and crushing are alternated between the two ore types, with crushed ore transported by conveyors approximately 3 km to the respective conical stockpiles located at the plants.

Challenges and successes at the Nkomati Nickel JV: pit to product process improvements Figure 5 clearly indicates the step change in total monthly milled tons for the complex, subsequent to the resolution of throughput constrains around the primary crusher in July/August 2012.

consistent hole depths as well as rigorous quality assurance inspections contributed to consistency in fragmentation as well as increased production.

MMZ comminution circuit

Ore fragmentation Fragmentation pattern improvements within the pit were realized through redesigned drilling patterns, and changes to the blast timing and blast direction. Interestingly, improved fragmentation was achieved at an increased hole spacing and much reduced powder factor. Production hole spacing was increased from 3.0 m × 3.5 m to 3.5 m × 3.5 m, while maintaining the 10.7 m hole depth and 3 m stemming material depth. Powder factor was reduced from 1.9 kg/m2 to 1.4 kg/m2.

Analysis of blasted material indicated that the majority of large rocks originated from the collar, close to the surface.

The introduction of 3 m stab holes (with a relatively light charge) between production holes resolved this problem.

Fragmentation was further improved through changing the direction of the blast from north to south (up-dip) to west to east (cross-dip) and introducing substantially slower blast timing. The introduction of hole depth counters to ensure

The MMZ circuit design (Figure 6) employs a FAG primary mill in closed circuit with a vibrating screen and pebble crushers. A secondary ball mill in closed circuit with cyclones supplies feed to the float circuit at SG 1.3–1.34, 70% -75 μm at 620 t/h. Design considerations and alternative comminution options considered were discussed by Wolmarans and Morgan (2009).

The choice of this circuit caused extensive debate, as the FAG mill was viewed by some as a ‘stone washer’, that would result in excessive metal losses through the sliming of softer nickel minerals. The proponents advocated that the reduced operating cost far outweighed potential recovery losses.

This circuit is very sensitive to feed particle size distribution as well as ore hardness. Excessive fine material (and coarse material) results in overloading of the mill.

Maintaining a full stockpile of 20–30 kt live capacity contributes greatly to mill stability through consistent feed.

Figure 5—Metallurgical complex milled throughput

Figure 6—MMZ milling circuit flow sheet with sampling and monitoring points

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Challenges and successes at the Nkomati Nickel JV: pit to product process improvements

MMZ FAG circuit pebble crushers Optimization of the MMZ milling circuit highlighted the critical importance of the gap setting of the pebble crushers.

While the process design criteria specified a closed side setting of 10–13 mm, realistically 19–21 mm was the best that could be achieved without causing frequent mechanical failures to the crushers.

Bypassing the pebble crushers or running with excessive gap setting would increase milling power from 24 to 30 kwh/t, with the mill feed rate cut back from 600 t/h to approximately 450 t/h due to overloading of the primary mill with pebbles. In addition, recoveries are affected by as much as 5–10%. While crusher comminution energy contributes just 400 kW to an average total of 16 MW for the FAG/ball mill/pebble crusher circuit (less than 2.5%), the pebble crusher is vital in removing critical-sized material from the circuit.

Extensive re-engineering of the crusher bushes and rigorous attention to planned maintenance, which was contracted to the OEM, allowed the crusher gap to be reduced gradually to 13 mm. The correct running in of the liners over a 4-week period was vital to prevent metal-on-metal contact and bush damage. Redesigned liners that do not need the extended running-in period are under development and are expected shortly.

A direct result of the higher-than-design crusher gap setting was an increased pebble crusher throughput, which exceeded the pebble production rate from the FAG mill. This resulted in stop/start operation of the crushers that threw ripples through the primary mill, mill discharge sump, and the cyclone circuits. Cyclical changes in froth stability were regularly observed in flotation, and were attributed to shifting grind as the milling circuit flows changed.

Data analysis indicated that stop/start operation of the crusher increases milling circuit power requirements by approximately 1 kWh/t, or R2.5 million per year at current power costs. Lost recovery costs are substantially higher (Van der Merwe, 2013).

Solving the crusher circuit instability was thus essential.

Various initiatives were tested to balance crusher rate feed with capacity. Although manipulating the screen panel size on the primary mill discharge was effective, knock-on effects on secondary mill performance resulted in less than optimum secondary mill performance.

A novel approach was to increase the crusher speed by approximately 15%. While it may sound counter-intuitive to increase a crusher speed to reduce capacity, this encourages choke conditions and reduced throughput, although the exact change in capacity has yet to be quantified. Further refinements have been to re-set pebble bin low and high levels to encourage more frequent (but shorter) stops and The Journal of The Southern African Institute of Mining and Metallurgy

starts. Downstream surge bins and other options are currently under consideration.

Milling circuit expert system tuning Tuning of the PxP mill expert control system and other control loops has assisted in improving stability. Assistance from the OEM, FL Smidth, was essential in understanding and optimizing the mill circuit control system in particular.

An important change in philosophy was the fixing of the number of operating cyclones and varying sump water addition, rather than allowing the control system to open and close cyclones to maintain constant pressure. The effects of varied density of the cyclone feed appears to be less disruptive than opening and closing cyclones.

MMZ comminution circuit operating costs While detailed modelling of the milling circuit is yet to be done, indications are that milling efficiency on the FAG/ball mill circuit is close to what would be expected in a crushing and ball milling circuit. Comminution circuit operating costs are lower on the MMZ plant than the PCMZ plant (which utilizes conventional crushing and two stages of ball milling) by approximately R10 per ton. MMZ plant recoveries are close to or in excess of design figures. The selection of this circuit design over others considered by DRA appears to have been justified.

MMZ flotation circuit The MMZ flotation circuit (Figure 7) is a relatively standard rougher-cleaner-recleaner configuration. A combined nickelcopper-cobalt and PGM concentrate is collected from the recleaner. Currently the pyrrhotite scavengers function as extended rougher capacity.

One of the immense challenges on the flotation circuit has been to achieve operational stability, and the resolution of a number of issues has contributed to the current relatively smooth operation.

Flotation cell level control Analysis of the cumulative level control valve outputs of the cells in the two parallel rougher banks indicated that the one bank received more flow than the other. In addition, the cell discharge valves would saturate at 100% output more regularly on the one bank, resulting in excessive rougher concentrate volumetric flow to the cleaner circuit. The unequal flow split was attributed to an inherent flaw in the design of the splitter between the two banks.

Excellent results were obtained with the implementation of the Gipronickel mass pull control algorithm. This relatively simple loop adjusts air to the flotation bank based on volumetric flow of concentrate. Although not as advanced as some concentrate mass pull control models, this system has been very effective in breaking the cyclical swings in recirculating load that characterized the circuit.

Flotation cell level control issues were exacerbated by the piping arrangement of the tailings from the rougher and scavenger banks. The gooseneck discharge pipes collect grit unable to be carried over with the tailings. Insufficient head on the gravity flow arrangement to the tailings thickener restricts further easy solutions, and alternative solutions are being considered.

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Various regimes of feeder operation have been attempted to supply consistent PSD material to the mill, with operation of one inner and one outer feeder generally used. On-line image analysis of mill feed to determine PSD, with automated feeder selection, has been less successful. Segregation patterns vary as the stockpile is loaded or depleted, and prediction of individual feeder PSD has not yet been successful.

Challenges and successes at the Nkomati Nickel JV: pit to product process improvements

Figure 7—MMZ flotation circuit flow diagram

Reagent optimization Talc and serpentine are the main contributors to high MgO levels in the concentrate, which incur severe smelter treatment penalties. Both these minerals have fast flotation kinetics and a strong froth stabilizing action. Underdepression results in vastly increased mass pulls, increased circulating loads, and low concentrate grades while also resulting in recovery losses.

Extensive tuning of the flotation circuit reagents has improved concentrate grades dramatically. Depressant allocation within the circuit has largely been shifted to the cleaners and re-cleaners. Careful control of collector has limited the over-collection of gangue.

The main value minerals recovered from the MMZ ore are pentlandite and chalcopyrite, while substantial (and variable) amounts of pyrite and pyrrhotite occur. Pyrite and chalcopyrite have substantially higher flotation kinetics than pentlandite, and mass pull control during periods of abnormal ore conditions is critical in maintaining recoveries.

Operator training It is an easy mistake for inexperienced operators to produce, for example, a high-grade pyrite concentrate at the expense of nickel recovery. While many operators had extensive exposure to PGM flotation, fewer had base metal flotation experience. Training of operators, control room operators, and metallurgists has paid off very well in terms of reducing and largely eliminating poor plant performance due to misdiagnosis or incorrect response to operational upset conditions. In combination with the ‘process recipes’ discussed below, operation of the MMZ plant has vastly improved.

Plant operational management

Performance mapping

Process recipes Of all the improvements implemented at the MMZ plant, the introduction of ‘process recipes’ has to rank among the most successful in ensuring consistent operation.

Process recipes stipulate the ranges within which key process variable must be run. These include froth depths, air

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addition rates, mill power draw, cyclone feed pressures, and SGs and other metallurgical variables. These have added immensely to the stability of the operation of the plant, and have prevented operator-induced instability due to incorrect or excessive corrections of perceived process upsets. Process recipes are adjusted infrequently, and only in consultation with senior management, metallurgists, and production staff and are signed off at senior level.

‘Short-interval control’ (SIC) was introduced as a way of guiding operators into checking the key process variables to assess the plant performance on a regular routine basis. This consists of a set of focused log sheets that required routine checks and monitoring, as well as monitoring of 2-hourly quality control assay results. These consist of a set of calculations attached to the 2-hourly quality control assay reporting sheets. Flotation circuit recoveries, upgrade ratios across flotation banks, ratios between Fe/MgO, Ni, and Fe, and other indicators of concentrate quality or ore quality warn operators of current of looming plant performance issues.

SIC is being extended with the introduction of time-instate (TIS) monitoring. TIS presents the operator with a ‘dial’ dedicated to each section of the plant (Figure 8). Data analysis has been used to assess the key parameters that influence the performance of a specific section, as well at the whole plant. An algorithm monitors multiple contributing factors, such as froth depths, aeration rate, concentrate sump levels, and flotation cell feed and product grades The operator is presented with and easy–to-read dial, with bar chart indicating what factor is out of range, and comments as to what ‘lever’ to pull to correct the situation.

Essential to the deployment of the TIS system has been the development and implementation of valid models that underlie the ‘idea state’ index.

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Extensive analysis of data using ‘performance mapping’ is used. Historical data pertaining to key variables considered by plant operational and metallurgical staff to be the main drivers and indicators of ideal operating performance is analysed, and the correlations are plotted in 2-dimensional space.

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Challenges and successes at the Nkomati Nickel JV: pit to product process improvements These performance maps have been extremely useful in indicating the correlations between various operational variables and process performance. They also indicate how often the process is in various ‘states’, which are often not the ‘ideal state’. The technique has been very useful in indicating appropriate process recipe set-points.

It is important to note that this method is used in conjunction with the plotting of trends and conventional metallurgical evaluation methods.

Graphics such as those shown in Figure 9 were generated for all major variables deemed to potentially impact plant performance, or to be indicators of conditions that would affect plant performance. The example shown correlates the rougher bank flotation performance with key parameters such as mill power, as well as other indicators such as the rougher motor power draw. While rougher cell power draw is not considered to be a key variable, it is an indicator of cell aeration, flotation feed densities, or wear on the flotation cell mechanism that requires scrutiny. These indicators are used together with the TIS dials to advise operators on what actions should be taken to rectify sub-optimum operating conditions.

Fluid Reports Fluid Reports is an automated reporting system developed by Blue Nickel in conjunction with U-Drive. This system draws data from the SCADA’s historian, and generates automated reports and trends on the performance of key process variables, including feed tons, key flow rates, power consumption, or similar process variables that directly affect throughput, recoveries, or product quality (Figure 10).

While not containing any information not already available on the SCADA, the system has proved extremely useful in producing easy-to-read trends tailored to an audience who do not have ready access to SCADA viewers.

An added benefit is the interpretation of trends relating to process control issues, inserted into the trends as ‘sticky notes’, and detailed in a separate report. Many of the metallurgical and control issues resolved were identified in Fluid Reports and resolved with the assistance of the Blue Nickel team.

A similar user-configurable web-based application allows real-time trends to be viewed from any computer or smartphone. These features are aimed at providing better

Figure 8—Time-in-state performance indication dials

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Figure 9—Performance maps

Challenges and successes at the Nkomati Nickel JV: pit to product process improvements

Figure 10—Fluid Reports trend reporting

support to operational staff outside office hours without the need for engineering, instrument, or metallurgical staff to be called out in the case of upset conditions.

Management systems Lastly, the benefits of strong focused leadership have to be recognized. Recent management changes, with fresh ideas and methods, have had a very positive influence on productivity.

Some of the most effective tools implemented were the ‘Gap’ list and ‘5 Why’ problem-solving methodologies. These contributed to a halving of the monthly mill trip rate in one month, from approximately 60 to less than 30 trips. Current efforts are aimed at dropping this to below 10 trips per month.

The importance of housekeeping on morale and discipline cannot be understated. The mills are arguably the largest, most expensive, and most visible items on a plant.

The condition and maintenance standards on the mills (Figure 11) provide immediate visual reference on the housekeeping and maintenance standards set for the rest of the operation.

must highlighted that this is not due to a failing of the management systems implemented, but rather the time taken to resolve systemic instrumentation issues.

Reduced mill breakdowns as well as the resolution of maintenance and process throughput issues at the primary crusher can largely be credited with the improved MMZ plant throughput, as can be seen in Figure 13. Step change improvements in milled tons can be seen from mid-July 2012.

Poor mill throughput in July 2012 was due to downtime attributed to pre-existing damage to FAG mill bearings and lube system. Throughput on the MMZ plant has exceeded nameplate tonnages regularly since, and is expected to do so more consistently on the conclusion of current improvement projects.

Plant performance improvements Numerous efforts were conducted simultaneously to resolve the metallurgical, operational, and management issues discussed above. Isolation of the contribution of individual process changes is thus difficult.

The impact of the introduction of new management techniques can probably best be seen in the reduction in mill trips. A step change in mill trips coincided with new management appointments in the middle of 2012 (Figure 12).

It can be seen that mill stoppages due to instrumentation issues persisted throughout the period ending June 2013. It

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Figure 11—Mills and lube rooms The Journal of The Southern African Institute of Mining and Metallurgy

Challenges and successes at the Nkomati Nickel JV: pit to product process improvements

Figure 12—MMZ mill breakdown analysis

Figure 13—MMZ plant thoughput

Conclusions Despite a number of challenges, the Nkomati Nickel JV has seen very strong growth in production over the last year.

A number of major equipment reliability and throughput issues at the mills and primary crusher were resolved, allowing for increased concentrate production. This has been achieved by optimization of the existing equipment, without major capital investment.

The successful introduction of problem-solving methodologies such Gap lists and 5 Whys played a significant role in identifying and addressing equipment and operational issues.

Close cooperation with OEMs was crucial in resolving the issues around both the Metso primary crusher and FLS pebble crushers.

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Improved fragmentation was achieved at lower powder factors, allowing improved primary crusher capacity and less downtime on the primary crusher.

Improved mill throughput and more stable milling circuit operation were achieved, with increased milling energy efficiency, through resolution of maintenance and operational issues around the pebble crushers. The importance of achieving design operational set-points on key equipment such as pebble crushers in a FAG circuit is clear. The operating cost of the FAG mill is approximately R10 per ton cheaper than the conventional crushing and ball milling circuit of the PCMZ plant.

Numerous flotation circuit optimization projects, including the implementation of process recipes, Gipronickel mass pull control, IME’s time-in-state monitoring, and performance mapping, as well as the use of Blue Nickel’s Fluid Reports, have collectively contributed to recovery improvements of 13.1% year–on-year (62.3% in 2011/2012 to 75.4% in 2012/2013).

Production trends indicate that the improvements are sustainable and that the figures for the year ahead should surpass the previous year’s performance.

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The metallurgical impact of the efforts to improve milling circuit stability, as well resolve the issues concerning the pebble crushers, are difficult to isolate from the impact of efforts to optimize the flotation circuit control.

While step change improvements in throughput are apparent the case of crusher and milled tons, gradual improvements in recovery are apparent. Figure 14 shows monthly nickel recovery figures.

Challenges and successes at the Nkomati Nickel JV: pit to product process improvements

Figure 14—MMZ plant nickel recovery

Mineralized Zone at Nkomati Mine. Internal memorandum, Norilsk Nickel Africa, 18 July 2008.

mining, engineering, production, and metallurgy. The input of various consultants, including the development of analytical and information systems tools, has been invaluable, as has been that of the engineering consultants and head office advisors.

VAN DER MERWE, K. 2012. Primary Crusher Report dd12Jul2012. Report, IME, 12 July 2012.

References

WOLMARANS, E. AND MORGAN, P. 2009. Milling circuit selection for Nkomati 375 ktpm concentrator. 5th Southern African Base Metals Conference, Kasane, Botswana, 27–31 July 2009. Southern African Institute of Mining and Metallurgy, Johannesburg. pp. 269–290. ◆

BRITZ, J. 2008. High level review on the mineralogical variability of the Main

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VAN DER MERWE, K. 2013. Pebble Crusher Performance March 2013. Report, IME, 13 March 2013.

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Evaluation of a Versatic 10 acid/Nicksyn™ synergistic system for the recovery of nickel and cobalt from a typical lateritic leach liquor

by A.C. du Preez* and M.H. Kotze*

Mintek has been involved in extensive test work since the early 1990s on the recovery of nickel and cobalt from leach liquors saturated in calcium, using synergistic solvent extraction systems. During this period the Nicksyn™ reagent was developed, optimized, commercially manufactured, and tested by Tati Nickel on a demonstration plant for more than 2800 operating hours. Efficient recovery of nickel without the co-extraction of calcium, thus avoiding gypsum formation in the extraction and stripping circuits, was illustrated. This synergistic system was recently evaluated on a laboratory scale for the recovery of nickel and cobalt from synthetic lateritic sulphate leach liquor containing about 3 g/L nickel, 0.5 g/L cobalt, 0.7 g/L manganese, 20 g/L magnesium, and with calcium at saturation.

Extraction and stripping parameters were determined for this feed liquor and are discussed in this paper.

Keywords solvent extraction, nickel laterites, reagents, synergistic systems.

Introduction The economic recovery of nickel from laterite ores has been pronounced for some years and will become more critical in the future, as lateritic ores constitute most of the world’s known nickel and cobalt resources, with nickel production from sulphide deposits progressively decreasing. High pressure acid leaching (HPAL) is being used for the recovery of nickel from nickel laterite ores, and increasingly atmospheric leaching is also being considered.

The quantity of laterite resources amenable to hydrometallurgical processing (limonite, nontronite/smectite) is almost twice that amenable to pyrometallurgical processing (saprolite, garnierite) (Bacon and Mihaylov, 2002).

Various hydrometallurgical flow sheets are being used for the recovery of nickel from laterite ores. Most plants use the Caron or HPAL processes, for which simple block flow diagrams are given in Figure 1. The Caron process is considered primarily for limonitic ores to avoid the high acid consumption associated with the iron content in the ore. The ore is calcined reductively to reduce the ferric (associated with goethite), prior to ammoniacal The Journal of The Southern African Institute of Mining and Metallurgy

* Mintek, Randburg, South Africa. © The Southern African Institute of Mining and Metallurgy, 2014. ISSN 2225-6253. This paper was first presented at the, Base Metals Conference 2013, 2–4 September 2013, Ingwenyama Conference & Sports Resort, Mpumalanga.

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Synopsis

leaching of the nickel and cobalt. Yabulu Nickel Refinery, Queensland, Australia implemented the Caron process.

The HPAL process can be used for limonitic as well as saprolitic ores (<4% Mg) and has been installed on numerous plants, including Moa Bay, Bulong, Murrin Murrin, and Goro (Bacon and Mihaylov, 2002).

However, the actual recovery of nickel and cobalt and their separation primarily from calcium, magnesium, and manganese is done employing different flow sheets. In the Murrin Murrin flow sheet the pH of the HPAL pregnant solution is adjusted to pH 3.5–4 to neutralize excess acid and precipitate most ferric, aluminium, and chrome. This is followed by sulphide precipitation of nickel and cobalt, which is the primary technology employed to separate the nickel and cobalt from manganese, magnesium, and calcium.

The Bulong flow sheet (Figure 2) also neutralized the free acid and precipitated ferric, aluminium and chromium, but it employed direct solvent extraction (SX) for the recovery of nickel and cobalt (Flett, 2005). Cyanex 272 (2,4,4-trimethylpentyl phosphinic acid) was used to recover cobalt from the neutralized stream, followed by nickel SX using Versatic 10 acid (V10, a tertiary-branched carboxylic acid). One of the major drawbacks of direct SX as operated at Bulong was that the selectivity of the V10 extractant was inadequate to prevent calcium loading during extraction.

This resulted in gypsum precipitation in the SX circuit, and hence major operational difficulties.

The most recent major nickel laterite project, namely Vale Inco’s Goro nickel project, was commissioned during 2012. Nickel is also recovered via direct SX (no prior precipitation

Evaluation of a Versatic 10 acid/Nicksyn™ synergistic system for the recovery of nickel (i)

Figure 1—Primary hydrometallurgical processing options for lateritic nickel ores (Dalvi, Bacon, and Osborne, 2004)

The very strong extraction of copper requires efficient removal of copper from the full flow of the pregnant leach solution via ion exchange (IX), which would be expensive. Furthermore, if any breakthrough from the IX circuit occurs, the copper would report to the Cyanex 301 circuit, requiring the copper to be stripped with thiourea in sulphuric acid medium (ii) Strong nickel and cobalt extraction makes stripping difficult. The Goro process was designed to use hydrochloric acid stripping in four stripping stages, each with 5 minutes’ residence time in the mixer, at an operating temperature of 60°C and a residual hydrochloric acid concentration of 3 M (iii) Due to the high residual acid concentration required for stripping, pyrohydrolysis is used for nickel recovery as NiO (iv) The introduction of chloride into the system requires more sophisticated materials of construction, and can have environmental concerns among other complications (v) Cyanex 301 is chemically unstable and in the presence of oxygen and metals such as ferric, the reagent is oxidized to form a disulphide. Hence, air has to be excluded from the operating system, which Goro achieves by employing Bateman Pulsed Columns. The reagent can be regenerated using sulphuric acid and zinc powder.

Mintek developed the Nicksyn™ reagent during the 1990s, and together with V10 it provides an alternative and more cost-effective approach to direct nickel and cobalt SX from laterite processing liquors. This synergistic system has been demonstrated over 2800 hours on the Tati Nickel Activox® demonstration plant in Botswana, and has since been commercialized (Du Preez et al., 2007; Masiiwa et al., 2008). This paper describes the technical performance of the V10 and Nicksyn™ synergistic system for the recovery of nickel and cobalt from neutralized HPAL lateritic leach liquor.

Figure 2—Simplified block flow diagram of the Bulong flow sheet

of nickel and cobalt) using Cyanex 301 [bis(2,4,4trimethylpentyl) dithiophosphinic acid]. Metal distribution equilibria generated for pH vs. extraction (Figure 3) show that nickel and cobalt can be extracted without neutralization during extraction (Mihaylov et al., 2000). However, this reagent has a number of drawbacks, namely:

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Figure 3—Metal extraction by 15 vol.% Cyanex 301 at an O:A phase ratio of 0.5 using NaOH for pH adjustment The Journal of The Southern African Institute of Mining and Metallurgy

Evaluation of a Versatic 10 acid/Nicksyn™ synergistic system for the recovery of nickel

Laboratory tests Analytical methods All metal analyses were done by inductively coupled plasmaoptical emission spectroscopy (ICP-OES) with a detection limit of 2 mg/L for all metals. Organic samples were stripped with sulphuric acid (approx. 1 M) at an organic-to-aqueous (O:A) phase ratio of 0.5, after which the strip liquors were submitted for analysis.

Reagents and solutions Versatic 10 acid (V10, a tertiary-branched carboxylic acid) was obtained from Chemquest (produced by Resolution Performance Products Ltd.), while Nicksyn™ was prepared for Mintek by an international, reputable manufacturer. The chemical composition and technical information on Nicksyn™ remain the proprietary information of Mintek and can therefore not be disclosed. Appropriate dilutions of V10 alone and V10 mixtures with Nicksyn™ were done using an aliphatic hydrocarbon diluent, Shellsol D70, which was obtained from Shell Chemicals.

Metal distribution studies Metal distribution equilibria (pH vs. extraction profiles, extraction and stripping isotherms) were generated by contacting the required organic phase with the appropriate feed solution at various O:A phase ratios, using rapid magnetic stirring and controlling the temperature in a waterjacketed glass vessel at 25ºC. An equilibrium time of 10 to 15 minutes was allowed to ensure steady state was reached.

The pH value of the aqueous phase (in the case of extraction isotherms) was adjusted or controlled by the addition of sodium hydroxide solution (approx. 1 to 10 M), using a calibrated combined glass reference electrode.

Samples of the organic phase were taken immediately after the aqueous samples to prevent possible re-equilibration after each pH adjustment. Aqueous samples were submitted for analyses. Organic samples were stripped with 1 M H2SO4 (O:A phase ratio of 0.5), after which the strip liquors were analysed for the relevant elements via ICP-OES. Organic phase (0.5 M V10 plus 0.25 M Nicksyn™) was batch-loaded for stripping purposes by contacting portions of fresh organic phase with synthetic laterite solution (at an O:A phase ratio of 0.45) at pH 6.0 and at 25ºC. Samples of aqueous and organic phases were analysed by ICP-OES. This procedure was repeated two more times to simulate the three stages required according to the McCabe-Thiele construction (Figure 7). The batch-loaded organic phase obtained was then contacted with a synthetic nickel spent electrolyte (approx. 71 g/L nickel in 40 g/L H2SO4) at different O:A phase ratios at 25ºC, measuring the final pH values of the loaded strip liquors. Samples of the loaded strip liquors and organic phases were analysed as previously described.

For the batch countercurrent extraction experiment, organic and aqueous phases were contacted (at 25°C), using magnetic stirring at an O:A phase ratio of 0.45. A sequence of batch contacts that simulates the conditions of a four-stage continuous flow process was used as illustrated in Figure 4.

Six full cycles (D to I, see Figure 4) were completed in this The Journal of The Southern African Institute of Mining and Metallurgy

way to ensure steady-state conditions. Samples of the raffinates of the fourth stages (4D to 4I), and the aqueous phases of the first (1I), second (2I), and third stage (3I) of the last cycle (I) were submitted for analyses. Portions of the loaded organic phases of the first stages (1D to 1I) as well as the loaded organic phases of the second (2I), third (3I) and fourth (4I) stages of the last cycle were taken and stripped as described above, after which the strip liquors were analysed

by ICP-OES.

Results and discussion Direct recovery of nickel and cobalt from a synthetic solution representing a nickel HPAL laterite leach liquor after iron, aluminium, and chromium removal was tested using the V10/Nicksyn™ system. Nicksyn™ is now commercially available, and hence offers a very attractive option to be considered for HPAL leach liquor.

Feed solution A synthetic feed solution was made up from metal sulphate salts to contain nickel, cobalt, manganese, calcium (at saturation), and magnesium. The average of various analyses of the synthetic laterite feed solution is given in Table I.

Figure 4—Scheme of contacts for batch countercurrent extraction experiment

Table I

Average composition of synthetic laterite feed solution Feed

Laterite leach liquor (synthetic)

Ni

Concentration, g/L Co Mn Mg

3

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Evaluation of a Versatic 10 acid/Nicksyn™ synergistic system for the recovery of nickel Organic phase compositions The different concentrations of V10 and the molar ratios of V10:Nicksyn™ diluted in Shellsol D70 are given in Table II.

Extraction metal-distribution equilibria (pH vs. extraction) The origin of the synergistic effect for nickel by a carboxylic acid (such as V10, which exists in the form of dimers H2A2), with the addition of a synergistic compound (L) such as Nicksyn™, has been discussed previously in terms of competing equilibria, and is given in Equations 1 to 3 (Du Preez et al., 2007; Masiiwa et al., 2008):

unaffected, hence the separation (ΔpH50Mn-Co) between cobalt and manganese increased from 0.75 to 1.13. Extractions of calcium and magnesium were negligible (<1%) under these conditions. This synergistic system therefore not only provides an option for the recovery and separation of nickel and cobalt efficiently from calcium and magnesium, but also gives the option of selecting a degree of manganese removal, with ease of pH control in practical flow sheets.

Extraction isotherms The distribution isotherms and McCabe-Thiele constructions for the extraction of nickel and cobalt from synthetic laterite leach solution generated using 0.5 M V10 plus 0.25 M

[1] [2] [3] where H2A2 denotes the carboxylic acid dimer and L denotes the synergist.

Results for selected pH vs. extraction isotherms are shown in Figure 5 and Figure 6. The pH50 values (the pH at which 50% of the metal originally present in the aqueous phase is extracted under a given set of conditions) are summarized in Table III. Synergistic shifts in the pH50 values for the extraction of nickel (i.e. the difference in pH50 value for V10 alone and the pH50 value for the appropriate synergistic mixture) increased from 1.23 to 1.68 units when Nicksyn™ addition was increased from 0.125 to 0.5 M, whilst the shifts for cobalt increased from 0.65 to 1.15 units with the same Nicksyn™ additions. The extraction of manganese was largely

Table II

V10 concentrations and molar ratios of V10:Nicksyn™ V10 Vol.% 9.6 9.6 9.6 9.6

Nicksyn™

V10:Nicksyn™

M

M

Molar ratio

0.50 0.50 0.50 0.50

0.125 0.25 0.50

4:1 2:1 1:1

Figure 5—Extraction of metals from synthetic laterite leach solution by 0.5 M V10 alone and 0.5 M V10 plus 0.125 M Nicksyn™ in Shellsol D70 at 25ºC

Table III

pH50 values for the extraction of metals from synthetic laterite leach solution using V10 alone and with Nicksyn™ in Shellsol D70 at 25ºC V10, M

Nicksyn™, M

0.5 0.5 0.5 0.5

0.125 0.25 0.5

pH50 Mg

Ca

Mn

Co

Ni

Mn-Co

Ca-Co

>7.17 >7.19 >7.43 >7.30

>7.17 >7.19 >7.43 >7.30

>7.17 6.62 6.47 6.50

6.52 5.87 5.53 5.37

6.37 5.14 4.97 4.69

>0.65 0.75 0.94 1.13

>0.65 >1.32 >1.90 >1.93

* A > sign preceding a pH value indicates that 50% metal extraction was not reached at this pH value

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Evaluation of a Versatic 10 acid/Nicksyn™ synergistic system for the recovery of nickel using a V10:Nicksyn™ ratio of 1:1 compared with the V10:Nicksyn™ ratio of 2:1, respectively, the additional costs of the increased Nicksyn™ concentration should be considered on an economic basis for each application.

In order to recover all cobalt together with nickel, the O:A phase ratio (and possibly the number of stages) would have to be adjusted to compensate for cobalt being ‘crowded off’ as illustrated in the unfavourable isotherms obtained for cobalt (Figure 9 and Figure 10).

The McCabe-Thiele construction redrawn for optimum cobalt recovery under the conditions tested indicated that a higher O:A phase ratio (1.86 vs. 0.45 as previously drawn for nickel recovery) should be employed to ensure a loading of approximately 257 mg/L cobalt without being ‘crowded off’ by nickel, using two countercurrent extraction stages. Under these conditions, nickel would still be recovered (leaving <2 mg/L in the raffinate) with minimum co-loading of calcium (5 mg/L), magnesium (4 mg/L), and manganese (7 mg/L).

For optimum cobalt recovery under these conditions the McCabe-Thiele construction indicated that an O:A phase ratio of 1.27 (instead of 0.33 previously drawn for nickel recovery)

Figure 6—Extraction of metals from synthetic laterite leach solution by 0.5 M V10 plus 0.25 and 0.5 M Nicksyn™ in Shellsol D70 at 25ºC

The Journal of The Southern African Institute of Mining and Metallurgy

Figure 7—Distribution isotherm for the extraction of nickel and cobalt from synthetic laterite leach solution by 0.5 M V10 plus 0.5 M Nicksyn™ in Shellsol D70 at 25ºC and pH 6.0

Figure 8—Distribution isotherm for the extraction of nickel and cobalt from synthetic laterite leach solution by 0.5 M V10 plus 0.5 M Nicksyn™ in Shellsol D70 at 25ºC and pH 5.8 VOLUME 114

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Nicksyn™ (at pH 6.0) and 0.5 M V10 plus 0.5 M Nicksyn™ (at pH 5.8) in Shellsol D70 at 25ºC are shown in Figure 7 and Figure 8, respectively. The McCabe-Thiele construction was drawn for optimum recovery of nickel with the idea to gauge what the possible recovery for cobalt could be under the chosen conditions. In order to determine the effect of insufficient aluminium removal prior to the SX circuit, about 200 mg/L aluminium (as sulphate) was added to the leach solution used for the generation of the extraction isotherms.

The McCabe-Thiele construction on the isotherm indicated that a loading of about 7.1 g/L nickel could be achieved in three countercurrent extraction stages at an O:A phase ratio of 0.45. The maximum loading of cobalt under these conditions was about 300 mg/L. Calcium and manganese co-extraction were about 6 mg/L and 34 mg/L, respectively.

The results for the McCabe-Thiele construction shown in Figure 8 indicated that a slightly higher loading of about 9.6 g/L nickel could be achieved with the increased ratio of V10:Nicksyn™ of 1:1 in three countercurrent extraction stages at an O:A phase ratio of 0.33. Cobalt loading under these conditions was slightly lower than 400 mg/L. The coextractions of calcium and manganese under these conditions were about 4 mg/L each.

Although slightly better separation (1.13 vs. 0.94 pH units) between cobalt and manganese, and higher loading of cobalt together with nickel (7.2 vs. 9.7 g/L) were achieved

Evaluation of a Versatic 10 acid/Nicksyn™ synergistic system for the recovery of nickel

Figure 9—Distribution isotherm for the extraction of cobalt from synthetic laterite leach solution by 0.5 M V10 plus 0.25 M Nicksyn™ in Shellsol D70 at 25ºC and pH 6.0

stages would be required, as shown in Figure 9 and Figure 10. A pH profile (and not a flat profile as was employed here) over all the stages could also assist to enhance cobalt recovery, providing no calcium is co-extracted.

The co-loading of some impurities at steady state of the batch countercurrent experiment is shown in Figure 12.

The efficient removal of aluminium prior to nickel and cobalt recovery is strongly indicated, as aluminium was strongly extracted by the synergistic mixture (from the feed solution containing approximately 213 mg/L only <2 mg/L was left in the raffinate). Magnesium extraction was low (<4 mg/L on the loaded organic for stages 1I, 2I, and 4I) and the anomaly observed in the higher loading in stage 2I (36 mg/L on the loaded organic phase) was most likely due to analytical error. Manganese was co-loaded (between 40 and 62 mg/L), with calcium co-loading minimal (approx.

7 mg/L) based on loaded organic phase analyses. Phase separations were clear in all stages and no crud formation was observed.

Further optimization of the O:A phase ratio, number of stages, and pH profile across the extraction circuit would have to be done to optimize cobalt recovery (including nickel) as well as limiting the co-loading of unwanted impurities such as manganese. This test work has to be performed for each individual lateritic feed solution.

Figure 10—Distribution isotherm for the extraction of cobalt from synthetic laterite leach solution by 0.5 M V10 plus 0.5 M Nicksyn™ in Shellsol D70 at 25ºC and pH 5.8

would be required to achieve a loading of about 400 mg/L cobalt without ‘crowding off’ by nickel in two to three countercurrent extraction stages. Under these conditions, a similar nickel recovery would still be expected (<2 mg/L in the raffinate) with minimum co-loading of calcium (4 mg/L), magnesium (4 mg/L), and manganese (14 mg/L). In both these cases optimum design should include economic cobalt recovery without calcium co-extraction.

Figure 11—Batch countercurrent extraction of nickel and cobalt from simulated laterite solution by 0.5 M V10 plus 0.25 M Nicksyn™ in Shellsol D70 in 4 stages and an O:A phase ratio of 0.45 and at 25ºC

Batch countercurrent test work A batch countercurrent test was performed for the extraction of cobalt and nickel with 0.5 M V10 plus 0.25 M Nicksyn™ in Shellsol D70 at 25ºC. Four extraction stages were used with an O:A phase ratio of 0.45 and an equilibrium pH value of 6.0 in each stage (i.e. a flat pH profile) over a total of six full cycles (see Figure 4). The results are illustrated in Figure 11.

A loaded organic phase (Stage 1I) containing about 7.7 g/L nickel was obtained at steady state (Cycle I, Figure 4), leaving about 5 mg/L nickel in the raffinate, which related to >99% extraction. Cobalt was loaded up to approximately 1.4 g/L in stage 3I, after which it was ‘crowded off’ by nickel to only about 176 mg/L on the loaded organic phase (Stage 1I) under these conditions. In order to recover both cobalt and nickel, a higher O:A phase ratio and probably more

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Figure 12—Batch countercurrent extraction of impurities from simulated laterite solution by 0.5 M V10 plus 0.25 M Nicksyn™ in Shellsol D70 in 4 stages and an O:A phase ratio of 0.45 and at 25ºC The Journal of The Southern African Institute of Mining and Metallurgy

Evaluation of a Versatic 10 acid/Nicksyn™ synergistic system for the recovery of nickel In order to conduct stripping test work, a batch of fresh organic phase (0.5 M V10 plus 0.25 M Nicksyn™ in Shellsol D70) was preloaded to reasonably represent a loaded organic phase that would be expected from an extraction circuit.

Synthetic spent nickel electrolyte was prepared to contain approximately 71 g/L nickel and 40 g/L H2SO4 (which should relate to a delta of approximately 24 g/L nickel). A stripping isotherm was generated and the results are shown in Figure 13.

Loaded strip liquor containing about 95 g/L nickel (at a pH value between 1.1 and 4.2) could be generated. The McCabe-Thiele construction indicated that two to three stages and an O:A phase ratio of 3.3 would be required to achieve this. Strip liquor pH values measured for stripping done at O:A phase ratios of 0.1 to 2.0 varied between 0.5 and 1.1, which indicates that adequate sulphuric acid was available for complete stripping of nickel. Stripping at O:A phase ratios of 5.0, 8.0 and 10.0 resulted in strip liquors exhibiting pH values of 4.2, 4.6, and 4.8 respectively, which indicated unfavourable stripping conditions for nickel. If plant operation required these operating conditions, it would be advisable to employ pH control (at about 3) in order to facilitate efficient stripping of nickel in a minimum number of stages and to provide an advanced electrolyte suitable for electrowinning.

Any co-loaded manganese could be removed by scrubbing with pH-adjusted water or a portion of the loaded strip liquor, depending on downstream requirements. The addition of a washing stage is recommended for removal of entrained aqueous phase from the loaded organic phase.

Conclusions ➤ The V10 plus Nicksyn™ synergistic mixture was evaluated for the use of direct solvent extraction of nickel and cobalt from lateritic leach liquor ➤ Nicksyn™ is now commercially available via a secured, reputable international supplier. It is a very attractive option and should be included in the evaluation for all HPAL hydrometallurgical projects ➤ For a solution containing about 3 g/L nickel, 0.5 g/L cobalt, 0.7 g/L manganese, 20 g/L magnesium, and

Figure 13—Distribution isotherm for the stripping of nickel from batchloaded 0.5 M V10 plus 0.25 M Nicksyn™ in Shellsol D70 with synthetic spent electrolyte at 25ºC The Journal of The Southern African Institute of Mining and Metallurgy

with calcium at saturation, the following parameters were found: – Three to four extraction stages would be required for optimum nickel and cobalt recovery. A pH profile (instead of the same pH value in every stage) could be employed to assist in the recovery of cobalt. The O:A phase ratio could vary and would be determined by optimum cobalt recovery excluding co-extraction of calcium – If required, a scrub stage could be included for removal of co-extracted impurities such as manganese and calcium. Typical scrub solutions would be pH-adjusted water, or a portion of the loaded strip liquor – One wash stage is recommended for removal of any entrained aqueous phase from the loaded organic phase – Two to three stripping stages would be required using spent electrolyte at an O:A phase ratio of about 3. The pH of stripping should be maintained at about 3 to facilitate a suitable advanced electrolyte for subsequent electrowining. ➤ Nicksyn™ was found to be stable over a testing period of 100 days (representing stripping conditions) in the laboratory. The solubility was found to be between 1 and 3 mg/L under specific conditions tested (Du Preez and Preston, 2004). ➤ Mintek is currently involved in an international project with a commercial client where Nicksyn™ will be included in the definitive feasibility study for nickel and cobalt recovery from laterite solutions.

Acknowledgements This paper is published by permission of Mintek. The authors would like to thank Nosipho Cola for her contribution to the generation of experimental data.

References BACON, G. and MIHAYLOV, I. 2002. Solvent extraction as an enabling technology in the nickel industry. Journal of the Southern African Institute of Mining and Metallurgy, November/December 2002. pp. 435–443.

DALVI, A.D., BACON, W., and OSBORNE, R. 2004. The past and the future of nickel laterites. PDAC 2004 International Convention, Trade Show & Investors Exchange, March 2004.

FLETT, D. 2005. Solvent extraction in hydrometallurgy: the role of organophosphorus extractants. Journal of Organometallic Chemistry, vol. 690. pp. 2426-2438.

MIHAYLOV, I., KRAUSE, E., COLTON, D.F., and OKITA, Y. 2000. The development of a novel hydrometallurgical process for nickel and cobalt recovery from Goro laterite ore. CIM Bulletin, vol. 93, no. 1041. pp. 124–130.

DU PREEZ, R., KOTZE, M., NEL, G., DONEGAN, S., and MASIIWA, H. 2007. Solvent extraction test work to evaluate a Versatic 10/Nicksyn™ synergistic system for nickel-calcium separation. Proceedings of the Fourth Southern African Conference on Base Metals, Swakopmund, Namibia, 23-27 July 2007. Southern African Institute of Mining and Metallurgy, Johannesburg. pp. 193–210.

MASIIWA, H., MATHE, O., DONEGAN, S., NEL, G., DU PREEZ, R., KOTZE, M., and IRELAND, N. 2008. Evaluation of Versatic 10 and Versatic 10/Nicksyn™ synergistic systems for nickel-calcium separation on the Tati BMR hydrometallurgical demonstration plant, Proceedings of ISEC 2008, Tucson, Arizona, September 2008. pp. 169-175.

DU PREEZ, A.C. and PRESTON, J.S. 2004. Separation of nickel and cobalt from calcium, magnesium and manganese by solvent extraction with synergistic mixtures of carboxylic acids. Journal of the South African Institute of Mining & Metallurgy, vol. 104, no. 6. pp. 333–338. ◆ VOLUME 114

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Evaluation of different adsorbents for copper removal from cobalt electrolyte

by V. Yahorava*, M. Kotze*, and D. Auerswald†

Synopsis Ion exchange is considered to be an effective technology for the removal of various impurities from cobalt advance electrolytes. With the correct choice of resin, ion exchange can consistently remove the required impurities to the levels for the production of high-grade cobalt metal. Although ion exchange was in the past used primarily for nickel removal, more recently it has been also considered for the removal of copper, zinc, and cadmium.

Generally, granular ion exchange products are used, but Mintek is currently evaluating ion exchange fibres for a number of applications, including the removal of copper from cobalt advance electrolytes. Fibrous ion exchangers have major advantages compared to granular resins in that they have significantly higher reaction rates, and wash water volumes could be limited.

Granular and fibrous ion exchangers were evaluated and compared for the removal of copper from cobalt advance electrolyte. A synthetic electrolyte containing 50 g/L cobalt and 50 mg/L copper was used for the test work. Equilibrium isotherms, mini-column tests, and split elution tests were done. The results were used to size a full-scale operation to treat 100 m3/h of electrolyte. The potential cobalt losses or recycle requirements were estimated, and data to calculate indicative operating costs for each adsorbent was generated. This information was used for a techno-economic comparison of granular and fibrous ion exchange systems for the removal of copper from cobalt advance electrolyte.

Keywords ion exchange, impurity removal, copper, cobalt electrolyte.

from one ionic form to the other without destruction of the filaments. The filtering layers of ion exchange fabrics do not significantly change their volume and permeability due to swelling of the filaments, which occurs as ion exchange proceeds and ionic state changes.

This paper demonstrates the possibility of using fibrous ion exchangers for hydrometallurgical applications such as copper removal from cobalt electrolyte. A comparison is presented of the design parameters and indicative costs for impurity removal using fibrous ion exchangers and granular resins having similar functionality.

Experimental procedures A number of IXFs were tested for their selectivity, maximum copper capacity, and equilibrium behaviour in order to choose the most suitable fibre for the subsequent investigations, which consisted of split elution, minicolumn, and mini-pilot plant tests. A similar test programme was carried out on a granular resin having the same functionality as the selected fibre, namely a resin containing imino-diacetic acid groups.

Maximum loading capacities for divalent metals Mintek conducted a significant amount of work on the development of ion exchange fibres (IXFs) during the 1980s and 1990s.

Scale-up of production of these materials was, however, not feasible at that stage. Similar development work was done at the Institute of Physical and Organic Chemistry (IPOC) in Belarus, who have commercialized the production of these materials more recently (Shunkevich et al., 2004; Soldatov et al., 2004; Vatutsina et al., 2007). This provided Mintek with the opportunity to resume its development on the use of IXFs for metallurgical processes.

An important feature of fibrous ion exchangers is their extremely high osmotic stability, which allows multiple cycles of drying and moistening as well as conversion The Journal of The Southern African Institute of Mining and Metallurgy

A maximum loading capacity test was performed in order to determine the maximum capacity of the fibres at different pH values for divalent metals using copper. A fibre-tosolution ratio was used that would provide an excess of metal in solution with respect to the

* Mintek, South Africa. † TENOVA Bateman Africa. © The Southern African Institute of Mining and Metallurgy, 2014. ISSN 2225-6253. This paper was first presented at the, Base Metals Conference 2013, 2–4 September 2013, Ingwenyama Conference & Sports Resort, Mpumalanga.

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Introduction

Evaluation of different adsorbents for copper removal from cobalt electrolyte amount of fibre used. This was done to ensure that there was adequate copper available, so that there were no equilibrium constraints during the determination of the maximum loading capacity. The tests were done by contacting air-dried fibre (in the H+ form) with the calculated volume of a 1.3 g/L copper feed solution at pH values of 2, 2.5, 3, 3.5, 4, and 4.5. Fibre samples were removed from the solution by filtration after the pH stabilized, washed with water, and stripped with an excess of 2 M HCl solution.

‘S-curves’ The effect of pH on the extraction of metals from an equimolar mixture of Cu2+, Mg2+, Mn2+, Ca2+, Ni2+, Zn2+, and Co2+ was determined to evaluate the relative affinity of each material tested for the metals of interest.

The mass of fibre used was sufficient (in capacity) to adsorb all the metals from solution. A feed solution was prepared from metal sulphate salts. One batch test was carried out for each fibre, and samples were taken at pH values of 2, 2.5, 3, 3.5, 4, 4.5, 5, 5.5, 6, 6.5, and 7. The contact period for each batch test was dictated by the time required for the pH to stabilize at the level at which it was controlled. The test was stopped when the pH had been stable for at least 1 hour. A solution sample was taken and the loaded fibre was filtered out of the barren solution. The loaded fibre was washed and stripped by contacting it with 300 mL of 2M HCl solution in a beaker while stirring using a magnetic stirrer for approximately 1 hour.

Similar experiments were done with the resin, but the feed solution did not contain Mn since it started to precipitate (as MnO2) in the presence of the resin at a pH of 3. The contact period for each pH point was 24 hours. The loaded resin was separated from the solution, washed, and eluted using 10 bed volumes (BVs) of 2 M H2SO4.

Equilibrium isotherms Adsorption equilibrium isotherms were generated for the loading of Cu and Co onto the adsorbents by batch contact of portions of the fibre/resin in the H+ form and synthetic solution (100 mg/L of Cu and 50 g/L of cobalt in a sulphate media) at different fibre to solution volume ratios. The pH values for the individual experiments were controlled at 3, 4, 4.5, and 5 for Fiban X-1 fibre, Lewatit TP 207 resin, Fiban K3, K-4, and AK-22 (3) fibres respectively.

Mini-column tests Breakthrough tests for TP 207 resin and X-1 fibre For the purpose of a techno-economic comparison between the FIBAN X-1 fibre and granular TP 207 resin, mini-column breakthrough tests were conducted employing the conditions presented in Table I.

Optimization of fibre column operating parameters Adsorption within a packed-bed column is a process in which continuous mass transfer occurs between two phases (the mobile phase containing the solution and the solid phase of the packed bed). The solution concentration in both phases is a function of the contact time and the depth of the adsorption zone. The mass balance of the packed-bed reactor can be described by the Thomas model (Thirunavukkarasu et al., 2002): [1] where: Ce C0 k q0

is the effluent adsorbate concentration (mg/L) the influent adsorbate concentration (mg/L) the Thomas rate constant (L/min·mg) the maximum solid phase concentration of solute (mg/g) m the mass of the adsorbent (g) V the throughput volume (mL) Q the volumetric flow rate (mL/min).

According to the mass balance of a specific packed bed, the determining factors of the mass balance for a given bed depth of the column are the volumetric flow rate and the initial solution concentration. For fibres it was also necessary to check the impact of fibre packing density because while this parameter changes, the filtering layer resistance also varies and this can influence the column efficiency.

Therefore, in order to optimize the adsorption process in a packed-bed fibre column, the following parameters were examined and their influence on the column efficiency was estimated: ➤ Flow rate, which translated to an increase in the linear velocity for the specific experimental set-up ➤ Packing density.

Efficiency of fibres in numerous operational cycles The objective of this experiment was to establish the efficiency of selected materials in numerous adsorption/elution cycles. A small column was packed with a fibre sample in the di-sodium form and synthetic solution simulating cobalt electrolyte containing 50 mg/L Cu as impurity was passed through the column to fully load the fibre. After loading, the fibre in the column was washed with excess water prior to elution with sulphuric acid. Eluate samples were analysed for their Co and Cu concentrations.

After elution, the fibre was converted back into the di-sodium form for the next adsorption cycle. These adsorption and elution cycles were repeated from 10 to 20 times.

After the last cycle, a portion of the stripped fibre was dried and analysed for any residual copper and cobalt.

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Table I

Mini-column test parameters Input

Lewatit TP 207

Material form

Fiban X-1

di-sodium

Density, g/cm3 (dry)

0.42

Absolute dry mass, g

119

0.23 5

Flow rate, mL/h BV/h

100 2

90 4

Aspect ratio of resin bed (height:diameter)

4.1

1.8

Volume of adsorbent (H+ form), mL

50

22

pH

4.5

Cu feed, mg/L

100

Co feed, g/L

50

The Journal of The Southern African Institute of Mining and Metallurgy

Evaluation of different adsorbents for copper removal from cobalt electrolyte While studying the influence of a certain parameter the values of the other parameters were kept constant. For all the tests 4 cm diameter columns were used. The tests were conducted on synthetic solution containing Cu at approximately 50 mg/L and Co at approximately 50 g/L. The feed solution pH was adjusted to a value of approximately 3 by the addition of 1 M NaOH solution. Input parameters for these tests are presented in Table II.

Mini-pilot plant The operational sequence for a lead-lag-lag configuration of columns is illustrated in Figure 1. Initially, all three columns are in adsorption. Transfer (loaded or lead resin/fibre column leaves adsorption circuit for elution) occurs when the bulk of the mass transfer zone (MTZ) has moved through the lead column (C1) into the lag columns (C2 and C3). The feed is transferred from the lead column (C1) to column C2, which then becomes the lead column. Column C1 is stripped and/or regenerated and returned to the adsorption circuit in the lag (or last) position. The solution is passed downflow through the resin bed, but upflow through the fibre bed. The main parameters of the process employing the fibre are presented in Table III. The entire operation could be divided into three parts: (1) Conversion of stripped fibre into the di-sodium form.

The dry X-1 staple was packed into the columns.

Regenerant (2 BVs of 1 M NaOH) was passed through the column at a flow rate of 10 BV/h (to convert fibre to the di-sodium form), followed by 1 L of wash water (upflow) to remove entrained base.

(2) Adsorption. Feed solution was passed through the adsorption column in an upflow direction. Once the fibre bed was filled with solution, the piston was pushed down to adjust the packing density to that required for the specific test (a sketch of the adsorption column with fibrous ion exchanger is presented in Figure 2). Solution was passed through at a flow rate of 4 L/h. Solution samples were taken

from the first and last columns every 15 minutes.

The first column was removed from the adsorption circuit to be eluted once the barren had reached 80% breakthrough, and the second column was transferred to the lead position.

(3) Elution. Air was pumped through the column to remove entrained solution from the fibre, then 200 mL of 0.2 M H2SO4 was passed downflow through the column at 14.2 mL/min (10 BV/h) to elute any Co loaded onto the fibre. After Co elution, 100 mL of 1 M H2SO4 was passed downflow through the fibre column at 14.2 mL/min (10 BV/h) to elute any remaining metals loaded on to the fibre. The stripped fibre was washed with 500 mL of water (upflow,

Figure 1—Illustration of the operating sequence of a lead-lag-lag ion exchange circuit

Table III

Parameters for fiber mini-pilot campaign Feed flow rate Cu loading Packing density Cu feed Number of columns Fibre volume per column Mass fibre per column

L/h

4

g/kg g/cm3 mg/L # mL g

25 0.4 50 3 83 33

Table II

Input parameters for optimization of the fibre column operation Material

Fiban X-1

Fibre form

Di-sodium

Parameter to be optimized BV/h mL g g/cm3

Parameter to be optimized Packing density Flow rate Fibre volume Mass of absolute dry fibre, H/D

11 95 26.6 0.28 2

16 95 25.7 0.27 2

22 95 25.7 0.27 2

34 95 25.7 0.27 2

44 95 25.7 0.27 2

Packing density* g/cm3 BV/h mL g

0.37 24 88 32.3 1.8

0.43 24 63 25.7 1.3

*Fibre with different capacity was used The Journal of The Southern African Institute of Mining and Metallurgy

0.53 26 60 32.1 1.2

Figure 2—Ion exchange fibre column for adsorption VOLUME 114

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Flow rate Fibre volume Mass of absolute dry fibre, Packing density H/D

Solution flow rate

Evaluation of different adsorbents for copper removal from cobalt electrolyte 50 mL/min) after elution. The piston compressing the fibre was lifted up and 200 L of 1 M NaOH was passed upflow through the column at 20 mL/min (the colour of the fibre became bright orange) to convert the fibre to the di-sodium form. The converted fibre was washed with 1 L of water to remove entrained NaOH (upflow, 67 mL/min). The entire elution and regeneration procedures of the fibre took 1 hour.

The column was returned to the system as the lag column in the third position, as shown in Figure 1. The fibre was compressed until the required volume was attained only after it was filled with solution. This was necessary because the fibre swelled in the base form and the resistance of the fibrous layer was high enough to block the solution flow through the column.

Results and discussion Characteristics of materials tested Four fibrous ion exchange materials (FIBAN®) were tested in comparison with a conventional granular ion exchanger. The imino-diacetic acid resin tested, Lewatit TP207, is a product from Lanxess. Functional groups and backbone structures of these materials are presented in Figure 3.

Although FIBAN® K-3 and K-4 have the same carboxylic functionality, their matrices are different. FIBAN® K-3 was synthesized from a polyacrylonitrile (PAN) backbone, while K-4 is a product of radiation grafting of acrylic acid to a polypropylene (PP) backbone (Shunkevich et al., 2004).

Generally, the difference in matrices of ion exchangers has an impact on their physical properties, and not their chemical properties. The main distinctive feature of the fibres synthesized on the PAN backbone is possible hydrolysis of the nitrile groups of the matrix over a period of time, resulting in the formation of additional carboxylic acid groups (Soldatov et al., 2004; Vatutsina et al., 2007).

predicted from the theoretical exchange capacities. This suggests that not all the functional groups within the structure of the fibres and the resin are available for the extraction of copper at lower pH values. Fiban AK-22(3) showed a higher capacity than was expected, which presumably was caused by complexation of copper by the polyamine groups on the fibre.

pH vs extraction isotherms The ‘S-curves’ (pH vs extraction) were constructed under conditions whereby the fibres/resins had sufficient capacity within one batch test in order to adsorb all the metals from solution (i.e. excess fibre/resin and limited metals in solution). The results of ‘S-curve’ tests can be used to establish the metal selectivity order, and the aim of this test work was to select the fibre that would provide the best copper selectivity over cobalt. The metal loading capacities increased with pH, and the S-curves achieved for Fiban X-1 are presented in Figure 4 as an example.

Results indicated that TP 207 and Fiban X-1would appear to function optimally at about pH 3 for Cu removal from Co electrolyte. The other fibres generally would require somewhat higher pH values. The order of selectivity for the various fibres and TP 207 were determined as follows:

Maximum copper loading capacities Maximum copper loading capacities as a function of pH for the various fibres indicated that the fibres function optimally at relatively high pH values. However, Cu could start precipitating at pH 5, so the maximum Cu loadings were done at pH 4.5. The results for the maximum copper loading capacities for the various fibres and the TP 207 resin tested are presented in Table IV. Loading capacities found experimentally were somewhat lower than the metal loading capacities that would have been

Figure 3—Chemical structures of materials tested

Table IV

Copper maximum loading capacity (pH 4.5) Adsorbent

Maximum theoretical capacity, meq/g*

Theoretical Cu maximum loading, mmol/g

Cu maximum loading, mmol/g

1.0 5.4 5.0 3.7 5.2

0.5 2.7 2.5 1.9 2.62

1.3 1.9 2.0 1.8 2.57

Fiban AK- 22(3) Fiban K-3 Fiban K-4 Fiban X-1** Lewatit TP207

*Based on content of –COOH groups **In some tests X-1 with 3.2 meq/g maximum theoretical capacity was used, namely, for the mini-pilot plant, re-usability testing and design

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Evaluation of different adsorbents for copper removal from cobalt electrolyte 22(3) were selected and subjected to numerous cycles of adsorption/elution to establish if any poisoning was evident.

The results of loading and elution cycles of Fiban X-1 and AK-22(3) are presented in Figures 6 and 7 respectively.

Analysis of residual Cu/Co at termination of the tests is listed in Table VI. Results indicate that Fiban AK-22(3) was poisoned with cobalt and completely lost its capacity within only 10 cycles of adsorption and elution. Hence, this material was excluded from further test work.

Fiban X-1 retained its selectivity and capacity over 20 cycles of adsorption and desorption (on average 0.82 mmol/g of divalent metals was loaded with only 4% variation in the Figure 4—Extraction versus pH for Fiban X-1

Fiban X-1: Cu > Ni > Zn Co > Mn > Ca > Mg Fiban K-3: Cu > Zn > Mn Ni Co > Ca > Mg Fiban K-4: Cu > Zn > Ni Co Ca > Mn > Mg Fiban AK-22(3): Cu > Ni > Co Zn > Mg Mn Ca Lewatit TP 207: Cu > Ni > Zn > Co > Ca > Mg The selectivity orders observed for Lewatit TP 207 and fibre Fiban X-1 were similar, which could be expected due to the similarity of their functional groups (Figure 3).

Equilibrium isotherms Equilibrium isotherms were constructed using synthetic cobalt electrolyte solution containing 100 mg/L Cu. The pH values for the individual experiments were controlled at 3, 4, 4.5, and 5 for Fiban X-1 and TP 207, K-3, K-4, and AK-22 (3) respectively. Results are presented in Figure 5.

The Langmuir equilibrium model (Soldatov et al., 2004) was fitted to the equilibrium data obtained and selectivity coefficients were calculated. The main results of the equilibrium tests are presented in Table V. Fiban AK-22(3) provided the highest loading of copper and the lowest co-loading of cobalt, while Fiban X-1, K-3, and K-4 had lower copper loadings and significantly higher Co co-loadings. Unnecessarily high Co co-loadings would generally increase the reagent amounts required to load and strip the fibre, and might cause a higher Co loss.

Langmuir parameters indicate that the resin had a higher maximum copper loading (the value of parameter a characterizes the saturation adsorption capacity). The highest affinity between the ion exchanger and copper was shown by AK-22(3) (b = 0.38 L/mg).

Based on these results the various adsorbents tested could be arranged in the following order with regard to their selectivity for copper over cobalt: Fiban AK-22(3) > Fiban X-1 > Lewatit TP207 > Fiban K-4 > Fiban K-3 Fibres with a selectivity for copper over cobalt higher than that of the Lewatit TP 207 resin were chosen for further test work.

Figure 5—Copper adsorption equilibrium isotherms

Table V

Main results of adsorption equilibrium tests Materials

X-1

K-3

K-4 AK-22(3) TP 207

Langmuir

a 26 47 39 b 0.04 0.01 0.01 RSQ 0.99 0.98 0.98

Maximum copper loadings Cobalt co-loading

mg/g

Selectivity Cu/Co*

20 10

32 0.38 0.92

164 0.01 0.95

24 72

21 55

31 13

97 82

1000 175

197

1304

591

where Curesin, Coresin - metal loaded onto the material; mg/g Cubarren, Cobarren – metal concentration in solution, mg/L.

Re-usability of Fiban X-1 and AK-22(3)

The Journal of The Southern African Institute of Mining and Metallurgy

Figure 6—Cobalt and copper loading of FIBAN X-1 over 20 adsorption/elution cycles based on elution data VOLUME 114

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Fiban AK-22(3) was noticed to change colour from white to pinkish, which indicated the possibility of cobalt poisoning of the fibre. Following this observation, Fiban X-1 and AK-

Evaluation of different adsorbents for copper removal from cobalt electrolyte Table VII

Mini-column breakthrough results

Cu loading, g/kg Co loading, g/kg Selectivity for Cu/Co Flowrate, BV/h Linear velocity, mm/sec Height of MTZ, cm

TP 207 resin

Fiban X-1 fibre

97 82 588 2a 0.057 8.2

35 28 638 4.7b 0.051 2.8

a Resin

volume was measured via tapped wet-volume method volume was controlled by pressing with a piston and could be varied depending on desired packing density

b Fibre

Figure 7—Cobalt and copper loading of FIBAN AK-22(3) over 10 adsorption/elution cycles based on elution data

Table VI

Co/Cu content in Fiban X-1 and AK-22(3) FIBAN®

Co, %

Cu, %

X-1 AK-22(3)

<0.05 2.26

<0.05 <0.05

loading capacity, average selectivity coefficient obtained was 996), and no residual copper/cobalt was found in the fibre after the final stripping cycle, as shown in Table VI.

Figure 8—Copper breakthrough curves for TP 207 and Fiban X-1

Mini-column tests

Mini-column breakthrough tests were done to allow a technoeconomic comparison between FIBAN X-1 fibre and the granular TP 207 resin. Synthetic solution containing 100 mg/L Cu and 50 g/L Co at pH 4.5 was passed downflow through the column containing fibre/resin at a flow rate of 2 BV/h. Results obtained are presented in Table VII and Figure 8.

Results indicated that the fibrous ion exchanger had a considerably shorter mass transfer zone compared to the resin, which would reduce the size of the plant significantly.

It also had a higher selectivity coefficient for Cu over Co, which should result in lower operating costs.

Influence of column parameters on performance Column parameters, including flow rate, packing density, and aspect ratio, were varied in order to optimize the design parameters for the fibre ion exchange column.

Fiban X-1 loading efficiency was tested at five different flow rates, and three different packing densities or aspect ratios. Results are depicted in Figure 9.

An increase in the linear flow rate increased the height of the MTZ, but decreased its residence time and therefore decreases the service time of the bed. However, the total

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BV till 1% breakthrough

Comparison of breakthrough curves for TP 207 resin and X-1 fibre

Packing density, g/cm3

Figure 9—Influence of the column parameters on bed volumes that can be treated for 1% Cu breakthrough: (a) flow rate at 0.27 g/cm3 fibre density; (b) packing density at 24 BV/h flow rate The Journal of The Southern African Institute of Mining and Metallurgy

Evaluation of different adsorbents for copper removal from cobalt electrolyte loading capacity of the fibre bed was constant (20 mg/g of Cu, 8 mg/g of Co). The choice of the optimum flow rate would then be dictated by design requirements.

Increasing of fibre packing density leads to an increase in the hydrodynamic resistance of the filtering layer to solution flow. At a packing density >0.4 g/cm3 a problem with the filtering layer resistance was observed. After conversion of the fibre into the basic form the fibre swells and blocks the free flow of solution at a high packing density. This is typical for carboxylic ion exchange fibres. Thus, the optimum packing density (taking into account the need to convert the fibre into basic form during current tests) was 0.4 g/cm3 for the laboratory equipment.

The following conclusions were drawn based on the results obtained during these tests where the parameters were varied: ➤ The influence of flow rate on fibre efficiency was described by exponential decay with variation of residence time for the mass transfer zone from 1.2 to 3 minutes ➤ The optimum packing density, where the pressure drop was reasonable, was 0.4 g/L. However, it was necessary to maintain a lower packing density during conversion of the stripped fibre into the basic form, rinsing of excess NaOH, and the initial adsorption phase. This was required as the resistance of the fibre bed hampered solution flow under these conditions.

The fibre should be pressed to a higher density only after adsorption has started.

Mini pilot-plant campaign results

➤ Somewhat better selectivity for copper over cobalt ➤ Shorter mass transfer height allows an increase in the productivity of the adsorbent ➤ It would allow a decrease in the total volume of fibre used per column; faster fibre regeneration also allows a decrease in fibre volume per column ➤ The total volume of the plant can be reduced by more than 90% using fibre, reducing the costs of the adsorbent as well as CAPEX ➤ Backwash might not be necessary ➤ Cobalt losses per annum are around 50% less for the fibre plant ➤ Various shapes and sizes of columns can be used.

CAPEX requirements The data presented in Table IX was used to size an ion exchange plant using either TP207 resin or the Fiban X-1 fibre as the adsorbent. An ’order-of-magnitude’ mechanical equipment cost was then calculated for each scenario.

The lower volume of fibre required results in considerably smaller columns, while the shorter cycle time reduces the volumes of all related tankage. The result is that the mechanical equipment cost required for the fibre plant is about 25% of that required for a resin-based plant.

Adsorbent cost can generally form a significant portion of the total capital cost of an ion exchange plant, especially where chelating resins such as the iminodiacetic acid resins are used. Because the adsorbent volume/mass is far lower for the fibre plant, further savings will be achieved in the cost of the ’first fill’.

The results of the countercurrent process are presented in Figure 10 and Table VIII. A lead-lag-lag fixed bed arrangement was used. Synthetic solution containing 50 mg/L Cu and 50 g/L Co at pH 3 was passed through the system at 4 L/h flow rate (total mass of absolute dry fibre was 94.5 g, total volume per 3 columns was around 240 mL).

As can be seen from the graph at the bottom of Figure 10, no copper breakthrough (> 0.5 mg/L in effluent stream) after the last column was observed during the mini-pilot plant tests. Ten transfers (of the lead columns) were done.

The average cobalt loss per column with respect to the delta between cobalt advance and spent electrolytes was 0.14±0.05 %. This could probably be lowered during optimization of the technique to selectively strip co-loaded Co prior to Cu stripping,, to achieve the levels that were obtained during the mini-column test work.

Comparison between fibre and resin plant designs

The Journal of The Southern African Institute of Mining and Metallurgy

Figure 10—Cu adsorption profiles (lead-lag-lag system)

Table VIII

Mini pilot-campaign results Average height of MTZ Average cobalt loss per column Average copper loading

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Based on the results of the mini pilot-plant campaign, and equilibrium and mini-column tests, the plant sizing, reagent consumption, and Co recycle/loss for TP207 resin and fibre Fiban X-1 were compared in order to establish the advantages or disadvantages of fibrous ion exchanger for hydrometallurgical applications. The results of calculations for plants using resin or fibre are presented in Table IX. In spite of the fact that the loading capacity of TP207 for copper is double that of the fibre, the use of Fiban X-1 has advantages compared with the resin:

Evaluation of different adsorbents for copper removal from cobalt electrolyte Table IX

Comparison of TP 207 resin with Fiban X-1 fibre Parameter

Fibre

Resin Input

Feed flow rate

m3/h

100

Cu loading

g/L

10

20

Co loading

g/L

10

45

min

1.6

Res time MTZ Cu in feed

mg/L

Co in feed

g/L

# columns

50000 2

Elution time

30 50

3

h

4

2

3

1.5

4

10

Output Upgrade

200 h

Fibre flow rate

m3/h

0.5

0.25

Fibre for MTZ

m3

2.7

50.0

Fibre for elution

m3

Fibre/column

m3

3.4

1.7

1.1

52.5

26.3

17.5

Total fibre volume

m3

6.8

5.1

4.6

105.0

78.8

70.0

1218

2436

3654

40

79

119

2363

1181

788

472.5

236.3

157.5

No. of transfers/year Time/year

6.8

400

Transfer time

3.4

2.3

210

0.75

105

70

2.5

h

8322

Co loss/recycle Co treated (based on 5 g/L delta) Co loaded per cycle

kg/a kg/cycle

4161 34

17

11

% Co loss/recycle per cycle (of the Co loaded)

kg/cycle

Co loss/recycle per annum

kg/a

Conclusions Several fibrous ion exchangers were investigated. Fiban X-1 with iminodiacetic acid groups proved to be the best adsorbent for copper removal from cobalt electrolyte. Test results obtained during evaluations of FIBAN® X-1 fibrous ion exchanger and granular ion exchanger Lewatit TP 207 were compared to assess the potential cost implications of a fibrous ion exchanger. This comparison established that the fibrous ion exchanger FIBAN® X-1 could be suitable for hydrometallurgical processes. Preliminary indications are that it could offer major cost savings compared to conventional granular iminodiacetic acid resin. Further work on Cu removal is focused on evaluation of the fibre and resin with limited or no regeneration included. Ultimately, the form in which the fibre or resin is to be employed will be an economic decision. Removal of entrained and co-loaded cobalt from the fibre is being optimized to further limit cobalt loss/recycle.

Other fibres investigated have promising characteristics for selective Zn and Ni removal from cobalt electrolyte.

FIBAN® X-1 could also be used as a substitute for TP 207,

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and for nickel-cobalt separation when these metals are present in solution in similar orders of concentration.

References SHUNKEVICH, A.A., MARTSINKEVICH, R.V., MEDYAK, G.V., FILANCHUK, L.P., and SOLDATOV, V.S. 2004. Comparison of fibrous carboxylic ion exchangers in water treatment to remove heavy metal ions. Russian Journal of Applied Chemistry, vol. 77, no. 2. pp. 249–253.

SOLDATOV, V., PAWłOWSKI, L., SHUNKEVICH, A., AND WASąG, H. 2004. New material and technologies for environmental engineering. Part 1. Syntheses and structures of ion exchange fibers. Drukarnia Liber Duo Kolor, Lublin.

127 pp.

THIRUNAVUKKARASU, O.S., VIRARAGHAVAN, T., SUBRAMANIAN, K.S., and TANJORE, S. 2002. Organic arsenic removal from drinking water. Urban Water, vol. 4. pp. 415–421.

VATUTSINA, V.M.., SOLDATOV, V.S., SOKOLOVA, V.I., JOHANN, J., BISSEN, M., and WEISSENBACHER, A. 2007. A new hybrid (polymer/inorganic) fibrous sorbent for arsenic removal from drinking water. Reactive and Functional Polymers, vol. 67, no. 3. pp. 184–201.

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The Journal of The Southern African Institute of Mining and Metallurgy

Thermodynamic analysis and experimental study of manganese ore alloy and dephosphorization in converter steelmaking

by G. Chen* and S. He*

In this study, the effects of slag compositions, slag amount, temperature, and carbon content of steel on the manganese and phosphorus distribution ratios during converter steelmaking were analysed using the classical regular solution theory, and industrial tests were performed using two 80 t top-and-bottom combined blown converters (duplex melting process). The results indicate that the slag amount, temperature, and carbon content in steel are the main factors affecting the manganese yield when converter slag compositions remain constant. The FeO content of the slag has a strong impact on the manganese distribution ratio, while the slag basicity and MgO content have no obvious effect. The calculations and experimental results show that the phosphorus distribution ratio increases sharply with increasing slag basicity R, but then decreases with the increase of MgO and MnO contents in the slag. The final slag in converter steelmaking should have the following characteristics: 3.5 < R < 4.5, 15% < (FeO) < 20%, and 6% < (MgO) < 8%. The slag amount should be controlled appropriately at the same time. The results of this investigation would be useful in deciding on the application of manganese ore in alloying and identifying the slagging regime in converter steelmaking.

Keywords slag compositions, distribution ratios, classical regular solution theory, slagging regime.

Introduction The increase in the international iron ore price in recent years has forced Chinese steel companies to seek domestic sources of iron ore in order to reduce costs. The availability of large amounts of phosphorus- and manganese-rich ores in the Three Gorges Reservoir region has made the development of a characteristic metallurgical technology in local steel plants possible (Wang and Dong, 2009). In the conventional steelmaking process, one converter must fulfill several functions, such as dephosphorization, decarburization, and raising temperature. It is unreasonable and uneconomical to decrease the phosphorus content using only one converter when high-phosphorus molten iron is smelted. Separation of the decarburization process from dephosphorization (using De-P and De-C converters) in the duplex melting process is advantageous, as it allows for the The Journal of The Southern African Institute of Mining and Metallurgy

Thermodynamic calculations The reaction of manganese ore during the alloying process in the converter is described as follows for a given slag system (Gao, Zhao, and Xing, 2011; Kaneko et al., 1993; Huang, X.H. 2008).

[1]

* College of Materials Science and Engineering, Chongqing University, Chongqing, China. Š The Southern African Institute of Mining and Metallurgy, 2014. ISSN 2225-6253. Paper received Mar. 2013; revised paper received Feb. 2014.

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Synopsis

use of phosphorus-rich iron ore and relatively lower amounts of slag, as well as direct alloying with manganese ore in the converter.

Consequently, iron procurement costs and metal losses are reduced to a considerable extent. For these advantages to be translated into commercial benefits, the dephosphorization converter should be fully exploited to produce low-phosphorus, high-carbon, hightemperature semi-steel, and the semi-steel smelting process should be further optimized by manganese ore alloying and by improving the manganese yield. A number of studies have been carried out on the improvement of the manganese yield and its impact on the converter, both in China and abroad (Suito and Inoue, 1995; Gao, Zhao, and Xing, 2011; Kaneko et al., 1993; Min and Fruehan, 1992; Lv et al., 2010; Soifer, 1958).

In this study, the main factors affecting the manganese yield are systematically investigated by thermodynamic analysis and industrial tests, and the relationship between manganese alloying and dephosphorization in the converter is discussed in detail. Finally, the final slag composition and control ranges for converter steelmaking are proposed.

Thermodynamic analysis and experimental study of manganese ore alloying [13]

[2]

[14] The equilibrium constant of the above reaction is [15]

[3] (Yang and Cao). The effect of [C] content of steel on the manganese distribution ratio (LMn = (%Mn)/[%Mn]) is given

by Equation [1] at PCO = 1, f[C] = 1, and f[Mn] = 1. The activity of (MnO), a[Mn], can be obtained by using the regular ionic solution model (Huang, 2008).

[4]

[5]

[6] [7] where x(i) is the mole fraction of positive ion i.

The (FeO) content plays a critical role in the control of manganese ore reduction during the later stage of blowing in the converter (Suito and Inoue, 1995; Gao, Zhao, and Xing, 2011; Morales and Fruehan, 1997; Takaoka et al., 1993).

This reaction is shown in Equation [8].

where the KP value is 0.0234 (Huang (2008). The values of x(P5+), x(Fe2+), γ(Fe2+), and γ(P5+) can also be obtained from the regular ionic solution model (Huang (2008).

Figures 1 and 2 show that R has no obvious effect on LMn. In addition, compared with (FeO), (MgO) has a less obvious effect on the change in LMn. LMn tends to increase with an increase in (FeO) but decrease with an increase in (MgO). Therefore, LMn is more strongly affected by the (FeO) content rather than R or (MgO) content. Figure 3 shows the variation of the calculated activity coefficient of (MnO) as a function of (FeO) content. When (FeO) content in slag increases from 15% to 35%, the activity coefficient of (MnO) decreases from 1.71 to 1.29, which can also be observed in the results of Jung et al. (1993), Jung (2003), and Suito and Inoue (1984). Obviously, the results obtained in this work are in agreement with the data from the literature. In addition, as can be seen in Figure 4, the activity of FeO in slag increases strongly with an increase of (FeO) in slag over the calculated concentration range, which is in good

[8] [9]

Equation [10] is derived from Huang (2008).

[10] The effects of slag basicity R = (%CaO)/(%SiO2) and (FeO) and (MgO) contents of the slag on LMn are studied through the reaction in Equation [8] at f[Mn] = 1. Similarly, the activity coefficients of (Fe2+), γ(Fe2+) and (Mn2+), γ(Mn2+) can be obtained from Equations [4] and [5] respectively.

Moreover, the effects of temperature on LMn are calculated separately by Equations [1] and [8], and the carbon content in Equation [1] is set as 0.08%.

The major dephosphorization reaction between molten steel and slag in the converter is described by Equation [11] (Basu, 2007; Ikeda and Matsuo, 1982).

Figure 1—Effect of FeO content of slag on LMn

[11] In order to calculate the phosphorus distribution ratio Lp = (%P)/[%P], the activity of the complex ion of phosphorous and oxygen can be expressed by the simplified reaction Huang (2008) : [12] Figure 2—Effect of MgO content of slag on LMn

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Thermodynamic analysis and experimental study of manganese ore alloying converter steelmaking are suited for dephosphorization. As shown in Figures 5–7, it is obvious that the dephosphorization effect increases sharply as R is increased, which is believed to dramatically reduce the activity coefficient of (P2O5) in slag, as recognized (Basu, 2007; Sobandi, Katayama, and Momon, 2002; Suito and Inoue, 1995; Turkdogan, 2000; Suito and Inoue, 1982, 1984; Nakamura, Tsukihashi, and San, 1993). A higher R is thus indispensable for dephosphorization in the converter.

Conversely, an excessively high R will worsen the kinetic conditions for manganese ore reduction and dephosphorization. Simultaneously, taking previous studies (Lv et al.,

Figure 3—Effect of FeO content of slag on r(MnO)

Figure 5—Effect of FeO content of slag on LP Figure 4—Effect of FeO content of slag on a(FeO)

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Figure 6—Effect of MnO content of slag on LP

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accordance with the results of other investigators (Morales and Fruehan, 1997; Huh and Jung, 1996; Sobandi, Katayama, and Momon, 2002). The distribution ratio LMn thus increases with the increase of (FeO), resulting in a low reduction efficiency of (MnO) when (the FeO) content is increased, as shown by Equation [9]. Hence, in order to improve the manganese yield in the converter, the (FeO) content of the final slag should be maintained at a low level in the alloying process performed using manganese ore. This has been affirmed by previous experiments (Suito and Inoue, 1995; Morales and Fruehan, 1997; Suito and Inoue, 1984; Jung, Rhee, and Min, 2002). However, the dephosphorization in the converter requires a high (FeO) content. As illustrated in Figure 5, the increase in (FeO) content initially enhances LP, but the trend is reversed beyond a certain level, and the optimal (FeO) content decreases with increasing R. Generally, LP has already reached the maximum level when the (FeO) content approaches 20% at a relatively high R value. These results are similar to those assessed thermodynamically and experimentally by previous researchers (Basu, 2007; Ikeda and Matsuo, 1982; Sobandi, Katayama, and Momon, 2002; Suito and Inoue, 1995). Therefore, the (FeO) content could be controlled between 15% and 20% to achieve a high manganese yield.

It is well known that slag with higher R and higher (FeO) content is required in the later stage of the smelting process for dephosphorization (Basu, 2007; Jeong et al., 2009; Nozaki et al., 1983). Consequently, the conditions used for

Thermodynamic analysis and experimental study of manganese ore alloying 2010; Basu, 2007; Tabata et al., 1990) into consideration, R should be high when the (FeO) content is relative low and, in general, the R of the final slag should be controlled between 3.5 and 4.5 to achieve a higher degree of dephosphorization.

Furthermore, LP clearly decreases with an increase in the (MgO) and (MnO) contents of the slag at a given R (Figures 6 and 7). Also, the influences of (MgO) and (MnO) contents on LP are gradually enhanced with the increase of R from 2 to 4; the effects of R on LP become progressively weaker with increasing (MgO) and (MnO) contents. Specifically, calculations show that increasing the (MgO) and (MnO) contents of the slag can increase the activity of (P2O5), which is detrimental for LP, as confirmed by Figure 8. Separately, the dephosphorization effects weaken significantly when the slag has a high (MnO) content, as has been reported by many investigators (Suito and Inoue, 1995; Mukherjee and Chatterjee, 1996; Simeonov and Sano, 1985) and observed by previous researchers, all of whom have advised against the addition of manganese oxide to converter slag. As a result, manganese ore alloying during converter smelting would become advantageous when low-phosphorus hot metal is used as raw material. Also, the experiments of Halder and co-workers demonstrated conclusively that 2CaO·SiO2 exists under conditions of higher slag basicity, lower steel tapping temperatures, and higher phosphorus contents of the hot metal, which comprised the majority of the solid part of the slag and also had greater solubility for phosphorus than the liquid part of the slag (Deo et al., 2004) . In addition, Suito, Inoue, and Takada (1981) also proved that slag containing 2CaO·SiO2 had a higher LP in the MgO-saturated slag of the system CaO-MgO-FeOx-SiO2, and the same result was also obtained in the CaO-CaF2-SiO2 system by Muraki, Fukushima, and Sano, (1985). However, the increase of (MgO) content in slag could result in a reduction in both the size of the 2CaO·SiO2 grains and the dissolution of phosphorus in 2CaO·SiO2. Thus, dephosphorization was greatly hindered (Deo et al., 2004). Therefore, taking only dephosphorization into consideration, the lower the MgO content the better. However, since (MgO) plays an important role in protecting the furnace lining, it is imperative that the appropriate amount of (MgO) should be present in the slag.

We can manipulate this relationship by using slag-splashing protection technology for the converter. In general, the (MgO) content should be controlled between 6% and 8%.

From Equations [1] and [8], the values of LMn are calculated separately as a function of steel temperature in Figure 9. The results of the authors and of Jung et al. (1993) and Jung, Rhee, and Min (2002) are plotted for comparison.

It is seen that most values of log(LMn) in this work are slightly higher compared with previous results. This may be a result of the slightly different components of the slag. In general, log(LMn) decreases linearly with increasing temperature. Equations [3] and [10] clearly indicate that the reaction in Equation [1] is endothermic and Equation [8] is exothermic. The equilibrium [Mn] content in steel is therefore expected to increase with increasing temperature. On the other hand, many studies indicate that the dissolution reaction (MnO(s) = MnO(slag) of (MnO) in slags is endothermic, so that (MnO) dissolution in slag increases with temperature (Jung et al., 1993; Suito and Inoue, 1984; Simeonov and Sano, 1985; Suito and Inoue, 1984.

Accordingly, the activity of (MnO) will increase with increasing mole fraction of (MnO) in the slag, and these results are widely accepted in previous studies (Ding and Eric, 2005). Thus, the equilibrium [Mn] content increases naturally.

Figure 10 shows that the final [C] carbon content of the steel has a great influence on LMn. With a decrease in the final carbon content [C] in steel, LMn increases distinctly, and this has been experimentally confirmed by previous investigators (Kaneko et al., 1993; Yang and Cao, 2009; Tabata et

Figure 8—Effect of MgO and MnO contents of slag on log(a(P2O5)

Figure 10—Effect of [C] of steel on LMn

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Figure 9—Effect of temperature on log(LMn)

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Thermodynamic analysis and experimental study of manganese ore alloying al., 1990; Matsuo, Fukagawa, and Ikeda, 1990). Clearly, the [C] in the molten steel can accelerate the reduction of (MnO) and improve the manganese yield. The [C] content of steel acts as a heat source and as well as a reducing agent, and carbon is thus consumed in large amounts. At the same time, carbon reacts with oxygen to form CO gas, which can enhance the fluidity of the entire slag system, which drives the reactions towards equilibrium (Keum et al., 2007).

The effect of the amount of slag on the manganese yield ((Ws * ([Mn]f - [Mn]s))/(0.3 * WMn), where [Mn]f and [Mn]s are the [Mn] content in the final steel and semi-steel respectively; Ws and WMn are the yield of final steel (tons) and manganese ore charged (tons), respectively) is shown in Figure 11. While the slag amount affects the manganese yield, it does not influence LMn. The manganese ore yield decreases sharply with an increase in the slag amount, due to the fact that the (MnO) content in the slag will decrease and the balanced [Mn] content in steel also will decline. As can be seen, since the manganese yield is less than 35% when the slag amount exceeds 60 kg/t, the slag amount must be maintained at a value less than 40 kg/t to ensure a high Mn yield (>45%). Similar results have been reported by other researchers (Kaneko et al., 1993; Tabata et al., 1990; Mukherjee and Chatterjee, 1996).

In conclusion, steelmaking by a process that involves manganese ore alloying and the use of low quantities of slag is an effective measure for lowering the consumption of raw materials and raising the manganese yield, which are the typical advantages of the De-P/De-C steelmaking process.

amount of manganese ore (0~810 kg) was added to ascertain the factors affecting alloying, which can provide reference data for further industrial-scale production.

Results and analysis The values of LMn obtained from the results are shown against (FeO) contents in Figure 12, where the R values range from 3.0 to 5.5. The values of LMn change irregularly

Figure 11—Effect of slag amount on manganese yield

Table II

Compositions of molten iron (mass %) Industrial tests C 4.0–4.2

Industrial process description To verify the results of the thermodynamic calculations, we carried out industrial tests using two 80 t converters at a steel plant in China. The blast furnace, De-P converter, De-C converter, refining, and continuous casting route has been been adopted. The experimental conditions used for the industrial-scale tests are shown in Table I. Tables II and III show the compositions of the molten iron and manganese ore used in the tests, respectively; the additions of manganese ore and compositions of semi-steel, final steel, and final slag are shown in Table IV. In this exploratory study, only a small

Si

Mn

P

S

0.4–0.6

0.2–0.4

0.11–0.45

0.036–0.06

Table III

Compositions of manganese ore (mass, %) CaO

MgO

SiO2

Al2O3

Mn

TFe

S

P

8–12

4–8

10–15

2–5

28–32

1–3

0.08–0.15

0.01–0.02

Table I

Operating conditions of De-P/ De-C converters De-P converter

De-C converter

Molten iron/semi-steel scrap

77~83 t 7~10 t Lime Slagging agent

84~93 t 0

Slagging elements

Returned slag Fluorite Bauxite

Slagging agent Manganese ore

10000–13000 Nm3/h

16000–16500 Nm3/h

300–600 Nm3/h

300–600 Nm3/h

7–9 min

10~12 min

Flow rate of bottom gas (N2/Ar) Blowing time

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Flow rate of top gas (O2)

Lime

Thermodynamic analysis and experimental study of manganese ore alloying Table IV

Manganese ore additions and compositions of semi-steel, final steel, and final slag (mass %) Semi-steel, %

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27

Mn Ore,

Final steel, %

Temp.,

Final slag, %

C

Mn

P

S

kg

C

Mn

P

S

°C

SiO2

Al2O3

CaO

MgO

TFe

P2O5

MnO

FeO

2.8 2.8 2.7 2.84 2.82 2.75 2.73 2.8 2.94 2.87 2.7 2.88 2.78 2.85 2.86 2.73 3.01 2.88 2.87 2.86 2.74 2.9 2.6 2.44 2.78 2.88 2.8

0.11 0.1 0.1 0.08 0.04 0.07 0.1 0.07 0.08 0.07 0.08 0.09 0.09 0.08 0.1 0.13 0.14 0.11 0.1 0.12 0.1 0.16 0.18 0.15 0.06 0.13 0.18

0.13 0.163 0.15 0.14 0.128 0.1 0.12 0.06 0.1 0.077 0.088 0.071 0.08 0.077 0.196 0.152 0.18 0.138 0.101 0.14 0.071 0.146 0.201 0.18 0.05 0.145 0.22

0.057 0.046 0.041 0.042 0.046 0.048 0.048 0.05 0.055 0.053 0.048 0.045 0.043 0.049 0.046 0.048 0.051 0.047 0.049 0.045 0.046 0.038 0.044 0.045 0.052 0.048 0.04

0 0 0 0 0 0 0 0 0 0 70 400 480 490 510 510 530 600 600 630 670 680 690 700 710 760 810

0.023 0.047 0.044 0.042 0.037 0.063 0.067 0.15 0.046 0.053 0.078 0.095 0.117 0.048 0.07 0.052 0.083 0.062 0.062 0.08 0.082 0.06 0.04 0.05 0.056 0.07 0.07

0.01 0.02 0.02 0.02 0.01 0.02 0.07 0.1 0.06 0.05 0.09 0.12 0.16 0.11 0.06 0.1 0.14 0.11 0.14 0.18 0.13 0.16 0.11 0.1 0.1 0.16 0.17

0.011 0.016 0.008 0.01 0.01 0.013 0.01 0.018 0.021 0.018 0.014 0.015 0.015 0.021 0.019 0.015 0.02 0.019 0.021 0.023 0.01 0.019 0.018 0.015 0.012 0.018 0.02

0.04 0.043 0.036 0.047 0.042 0.048 0.045 0.06 0.057 0.052 0.048 0.048 0.048 0.058 0.039 0.039 0.044 0.051 0.058 0.044 0.038 0.04 0.038 0.038 0.048 0.04 0.037

1612 1638 1608 1624 1630 1627 1638 1666 1642 1645 1632 1629 1640 1646 1624 1655 1653 1664 1634 1645 1622 1629 1653 1630 1621 1656 1660

8.34 16.9 8.12 8.08 8.51 9.52 8.98 12.2 10.58 10.02 10.24 10.78 11.24 12.62 8.84 8.12 9.38 8.14 11.7 9.86 10.46 9.12 8.42 8 9.86 9.46 8.5

2.05 2.57 2.57 2.05 1.54 2.05 2.05 1.94 2.62 3.05 3.19 2.9 3.2 3.07 1.05 2.05 2.57 2.57 2.88 1 2.64 1 1.8 1.05 2.01 1.04 1.04

40.66 46.09 43.27 44.82 43.41 42.15 38.34 38.76 44.26 40.74 41.72 41.16 40.45 43.7 41.3 38.48 40.88 34.39 42.01 42.43 40.45 38.9 39.33 40.6 37.49 38.9 42.15

8.61 8.01 7.4 6.18 6.08 7.9 8.72 12.67 8.82 11.15 10.34 9.12 8.82 11.25 10.95 7.9 8.41 8.92 8.82 7.8 9.12 10.54 10.84 8.82 9.93 7.09 8.61

24.02 18.43 22.9 21.5 23.32 20.94 23.04 20.25 16.34 18.71 18.29 17.03 17.87 14.92 22.06 22.06 17.59 24.02 17.45 17.45 18.71 18.29 19.97 23.04 21.78 18.99 17.87

2.87 3.61 3.34 3.9 3.52 3.57 3.55 2.52 1.76 1.76 1.51 1.26 1.76 2.01 2.89 3.58 4.07 3.57 2.01 2.18 1.51 4.04 3.43 3.46 1.26 4.13 4.41

2.22 2.22 1.93 1.93 2.22 1.93 3.86 2.53 2.97 2.82 4.46 5.94 6.54 4.31 3.4 4.75 4.75 5.2 5.5 6.06 6.13 6.65 5.02 5.62 6.09 6.21 6.35

21.7 16.88 21.2 20.19 21.84 20.19 20.98 19.26 17.17 18.75 17.53 17.68 17.89 14.37 20.55 21.77 17.24 22.42 16.96 15.45 19.4 16.31 16.81 19.76 21.12 15.59 15.59

with the increase of R at a given (FeO) content, and LMn increases with an increase in (FeO) content at a given R, which shows a relatively consistent tendency with thermodynamics (Figure 1). This is also confirmed in Figures 13 and 14. LMn increases as the FeO content increases when [C] content and temperature remain constant. In addition, LMn shows a rapid increase as the [C] in the steel decreases at the same FeO content (Figure 13), and the temperature has a similar impact on LMn, as illustrated in Figure 14. In Figure 13, the experimental values of LMn locate between 25 and 45, indicating that these values agree well with the calculated values (which are between 20 and 47) according to Equation [1] (Figure 10), under the condition of 0.04–0.09% [C] carbon content. However, these experimental values turn out to be somewhat less than the calculated values according to Equation [8] (Figure 10), probably because the reduction of MnO in slag by carbon [C] in steel leads to a significant increase of the [Mn] content of the steel, and in turn, the oxidization of [Mn] by FeO in the slag may not reach a thermodynamic equilibrium. As shown in Figures 12–14, as expected, low LMn values result when the (FeO) content is less than 25%, the [C] content exceeds 0.06%, and the temperature is higher than 1920K. As is well known, the increase in temperature depends on the oxidization of carbon during blowing in the BOF, and there is a strong negative correlation between [C] carbon content and temperature. On the other hand, the oxidation of [C] carbon will be depressed while the oxidation of [Fe] will be facilitated when the [C] content is less than about 0.05% (Huang, 2008), which will cause excessive (FeO) content in slag. As demonstrated in Figure 15, when the final carbon content is less than 0.06%, the majority of the corresponding (FeO) content in slag

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Figure 12—Effect of FeO content of slag on LMn

Figure 13—Effect of [C] on LMn The Journal of The Southern African Institute of Mining and Metallurgy

Thermodynamic analysis and experimental study of manganese ore alloying tests are fairly low (usually below 20%); this is the inevitable consequence of the presence of a large amount of slag. The main reason is that the iron ore used by the aforementioned steel plant is rich in phosphorus, and the [P] content of the semi-steel charged in the De-C furnace is still high and a large amount of time and energy will lost in dephosphorization. Consequently, a large amount of slag is generated.

At the same time, a large amount of slag is required for deep dephosphorization, which in turn has an enormous negative influence on manganese yield.

The influences of R and the (FeO), (MnO), and (MgO) contents of the slag on LP are shown in Figures 17–19. From the experimental results, it is obvious that LP increases with Figure 14—Effect of temperature on LMn

Figure 17—Effect of FeO content on LP Figure 15—Relation between of FeO content and slag amount

Figure 18—Effect of MnO content on LP Figure 16—Effect of slag amount on manganese yield

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Figure 19—Effect of MgO content on LP VOLUME 114

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exceeds 28%. However, a high [C] carbon content, high temperature, and low (FeO) content are required in order to attain a low LMn. More intensive research is required to determine the appropriate amount of carbon powder that should be added to the slag when the [C] carbon content of the semi-steel is insufficient to provide adequate heat and reduce the manganese ore to a large extent.

The influence of the slag amount on the manganese yield is shown in Figure 16. The yield of the manganese ore decreases sharply with an increasing slag amount at similar values of LMn, and it falls below 25% when the slag amount is 60 kg/t. Generally, the manganese yields in the industrial

Thermodynamic analysis and experimental study of manganese ore alloying increasing R from 3.0 to 5.5, regardless of whether the (FeO), (MnO), or (MgO) contents are held constant. It is also found that the higher the (FeO) content of slag, the more the LP increase at a given R. However, the optimum (FeO) content is not observed. This is corroborated by the theoretical calculations and has been reported in previous investigations (Basu, 2007; Ikeda and Matsuo, 1982; Sobandi, Katayama, and Momon, 2002; Suito and Inoue, 1995). It is considered that this phenomenon is due to the increase in the amount of slag, even if the LP is decreased slightly with an (FeO) content higher than the optimum. As shown in Figure 15, the results confirm a strong positive correlation between (FeO) content and slag amount. In addition, as expected, the dephosphorization effect weakens markedly with increasing (MgO) and (MnO) contents in slag when R is greater than 4.

However, when R is less than 4, no regular relationship has been found. Furthermore, the impact of R on LP is weakened when the (MgO) and (MnO) contents are high, respectively.

From the data pertaining to the smelting test conducted in 27 heats, we can conclude that the average final [P] content of the steel and the average degree of dephosphorization are 0.016% (0.008–0.023%), and 87.4%, respectively, when the average [P] content of the semi-steel charged in the De-C furnace is 0.126% (0.05–0.22%). Since a larger amount of slag (on average, about 68.3 kg/t) with high (FeO) content (mean, 25.6%) is used in the converter to decrease the [P] content to the required steel grade, a relatively good degree of dephosphorization is achieved, but the manganese yield is only 17.2% on average. Hence, for manganese ore alloying to be beneficial, the process in the De-P furnace should be optimized to decrease the [P] content of the semi-steel further.

Obviously, the data obtained in the test work is basically in good agreement with the results of the thermodynamic calculations. This, in turn, shows that the choice of the method of calculation is reasonable.

On the basis of the thermodynamic calculations and the industrial test results, it is concluded that coordinated control between the dephosphorization ability and manganese ore alloying technology in the De-C converter should be considered carefully. The characteristics of the final slag for converter steelmaking should be controlled in the following ranges: 3.5 < R < 4.5, 15% < (FeO) <20%, and 6% < (MgO) < 8%.

Conclusions A thermodynamic analysis and industrial tests of manganese ore alloying and dephosphorization in converter steelmaking were carried out. The conclusions can be drawn as follows.

(1) The main factors affecting the alloying process performed using manganese ore in the converter are the slag amount, temperature, and the [C] content of the steel with a given slag system (2) The (FeO) content of the slag has an enormous impact on LMn but shows no clear relationship with the slag basicity or the (MgO) content of the slag (3) The LP increases sharply with increasing slag basicity, but weakens with increasing (MgO) and (MnO) contents in the slag (4) The characteristics of the final slag for converter steelmaking should be controlled in the following

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ranges: 3.5 < R < 4.5, 15% < (FeO) < 20%, and 6% < (MgO) < 8%. The slag amount should be controlled appropriately at the same time.

Acknowledgements The authors greatly appreciate the funding support from Chongqing Science and Technology Key Project (CSTC2008AB4018)

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MORALES, A.T. and FRUEHAN, R.J. 1997. Thermodynamics of MnO, FeO, and phosphorus in steelmaking slags with high MnO contents. Metallurgical and Materials Transactions B, vol. 28B. pp. 1111–1118.

MUKHERJEE, T. and CHATTERJEE, A. 1996. Production of low phosphorus steels from high phosphorus Indian hot metal: experience at Tata Steel. Bulletin of Materials Science, vol. 19. pp. 893–903.

MURAKI, M., FUKUSHIMA, H., and SANO, N. 1985. Phosphorus distribution between CaO-CaF2-SiO2 melts and carbon-saturated iron. Transactions of the Iron and Steel Institute of Japan, vol. 25. pp. 1025–1030.

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Thermodynamic analysis and experimental study of manganese ore alloying NAKAMURA, S., TSUKIHASHI, F., and SAN, N. 1993. Phosphorus partition between CaOsatd.-BaO-SiO2-FetO slags and liquid iron at 1873K. ISIJ International, vol. 33. pp. 53–58.

SUITO, H. and INOUE, R. 1995. Thermodynamic assessment of hot metal and steel dephosphorization with MnO-containing BOF slags. ISIJ International, vol. 35. pp. 258–265.

NOZAKI, T., TAKEUCHI, S., HAIDA, O., EMI, T., MORISHITA, H., and SUDO, F. 1983.

Mechanism of hot metal dephosphorization by injecting lime base fluxes with oxygen into bottom blown converter. Transactions of the Iron and Steel Institute of Japan, vol. 23. pp. 513–521.

SUITO, H. and INOUE, R. 1995. Thermodynamic assessment of manganese distribution in hot metal and steel. ISIJ International, vol. 35. pp. 266–271.

SIMEONOV, S.R. and SANO, N. 1985. Phosphorus equilibrium distribution between slags containing MnO, BaO and Na2O and carbon-saturated iron for hot metal pretreatment. Transactions of the Iron and Steel Institute of Japan, vol. 25. pp. 1031–1035.

SOBANDI, A., KATAYAMA, H.G., and MOMON, T. 2002. Activity of phosphorus oxide in CaO-MnO-SiO2-PO2.5(-MgO, FetO) slags. ISIJ International, vol. 38. pp.781–788.

SOIFER, V.M. 1958. Use of manganese qre in scrap-process steelmaking.

Metallurgist, vol. 2. pp. 170–172.

Suito, H. and Inoue, R. 1984. Manganese equilibrium between molten iron and MgO-saturated CaO-FeOt-SiO2-MnO slags. Transactions of the Iron and Steel Institute of Japan, vol. 24. pp. 257–265.

SUITO, H. and INOUE, R. 1984. Phosphorus distribution between MgO-saturated CaO-FetO-SiO2-PO5-MnO2 slags and liquid iron. Transactions of the Iron and Steel Institute of Japan, vol. 24. pp. 40–46.

Suito, H. and Inoue, R. 1984. Thermodynamic considerations on manganese equilibria between liquid iron and FetO-MnO-MOx (MOx=PO2.5, SiO2, AlO1.5, MgO, CaO) slags. Transactions of the Iron and Steel Institute of Japan, vol. 24. pp. 301–307.

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SUITO, H., INOUE, R., and TAKADA, M. 1981. Phosphorus distribution between liquid iron and MgO saturated slags of the system CaO-MgO-FeOx-SiO2.

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TABATA, Y., TERADA, O., HASEGAWA, T., KIKUCHI, Y., KAWAI, Y., and MURAKI, Y. 1990. Manganese partition equilibrium in less slag blowing at BOF linked to high speed dephosphorization of hot metal. Tetsu to Hagane-Journal of the Iron and Steel Institute of Japan, vol. 76. pp.1916–1923.

TAKAOKA, T., SUMI, I., KIKUCHI, Y., and KAWAI, Y. 1993. Manganese reaction rate in combined blowing converter with less slag. ISIJ International, vol. 33. pp. 98–103.

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Modern Management Science. pp. 72–73.

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SIMEONOV, S.R. and SANO, N. 1985. Manganese equilibrium distribution between carbon-saturated iron melts and lime based slags containing MnO, BaO, and Na2O. Transactions of the Iron and Steel Institute of Japan, vol. 25. pp. 1116–1121.

SUITO, H., and INOUE, R. 1982. Effect of calcium fluoride on phosphorus distribution between MgO saturated slags of the system CaO-MgO-FeOx-SiO2 and liquid iron. Transactions of the Iron and Steel Institute of Japan, vol. 22. pp. 869–877.

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Atmospheric oxidative and non-oxidative leaching of Ni-Cu matte by acidified ferric chloride solution

by L.M. Sekhukhune*, F. Ntuli*, and E. Muzenda*

The atmospheric leaching of copper-bearing matte by acidic ferric chloride solution was studied at the laboratory scale. The aim was to achieve maximum copper and nickel recovery by investigating the mechanisms of leaching, as well as identifying the effect of temperature, and concentration of ferric chloride and oxygen. Djurleite (Cu1.96S), hazelwoodite (Ni3S2), and Ni alloy were the primary phases detected in the matte. The quantitative composition of the matte was Cu 31%, Ni 50%, S 13%. Fe and Co constituted 2%, with platinum group metals (PGMs) accounting for 0.5%. A maximum nickel extraction of 98% was achieved using two-stage oxidative leaching at 90°C and 11 g/L Fe3+ as compared to 65% under non-oxidative conditions. A copper extraction of 99% was achieved in the first 45 minutes using two-stage non-oxidative leaching, and copper was recovered from solution by cementation. Three processes took place simultaneously throughout the leaching process, namely: dissolution, cementation/ metathesis, and oxidation. The leaching process was found to be diffusioncontrolled.

Keywords Ni-Cu matte, acid leaching, cementation, ferric chloride, leaching mechanism.

Introduction Hydrometallurgy has become one of the most economical processes for recovering metals from low-grade matte. Matte is a mixture of metal sulphides and precious metals produced from the smelting of sulphide ores. Acid leaching unselectively separates the base metals from the precious metals, producing a liquor rich in base metals and leaving a solid residue rich in precious metals. Solvent extraction (SX) and electrowinning (EW) can be used to recover marketable base metal products from the liquor. Gaseous reduction can also be employed to produce base metal powders (Agrawal et al., 2006).

There are different types of leaching techniques and each type has requirements in terms of particle size, leaching time, agitation rate, etc. to ensure an efficient process and high percentage metal extraction. Agitation leaching in batch vessels operated under atmospheric or high pressure has been widely used commercially for processing Ni, Cu, and Co sulphide concentrates. Leaching of these The Journal of The Southern African Institute of Mining and Metallurgy

* Department of Chemical Engineering, University of Johannesburg, Johannesburg, South Africa. © The Southern African Institute of Mining and Metallurgy, 2014. ISSN 2225-6253. Paper received Jun. 2013; revised paper received Nov. 2013.

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Synopsis

mattes has been conducted using ammonia or acid solutions following the developmental work of Sherritt Gordon in the period 1950–1969 (Forward and Mackiw, 1955; Pearce et al., 1960). The choice between acid and ammonia leaching has been based on local conditions or the composition of the matte.

Acid leaching has been used for mattes with substantial cobalt contents (above 3%), while ammonia leaching has been widely used for Ni-Cu mattes with low cobalt contents (Pearce et al., 1960). However, with further developments in acid pressure leaching technology by Sherritt Gordon in the 1960s most Ni-Cu mattes have been treated using acid leaching. Impala Platinum commissioned the first commercial acid pressure leaching process for treating Ni-Cu mattes containing platinum group metals (PGMs) in horizontal autoclaves in 1969, and since then acid leaching has been widely used in the South African platinum industry (Plasket and Romanchuk, 1978).

Acid leaching can be conducted in a sulphate or chloride medium. The sulphate medium has been widely used, especially in South Africa, because the process equipment used is more adaptable to sulphate rather than chloride systems and there is no possibility of platinum dissolution as in chloride systems (Brugman and Kerfoot, 1986). However, chloride systems have the following advantages; most metal chlorides are more soluble than sulphate salts, leaching can be performed at moderate temperatures, and the oxidation process yields elemental sulphur, which is environmentally more acceptable than sulphate from sulphuric acid leaching (Park et al., 2006). Given these advantages and the development of corrosion-resistant material,

Atmospheric oxidative and non-oxidative leaching of Ni-Cu matte the leaching behaviour of typical South African PGMcontaining Ni-Cu mattes in chloride media is worth investigating.

This study investigated the atmospheric leaching of PGMcontaining Ni-Cu matte by acidic ferric chloride solution. The aim of this project was to achieve maximum Ni and Cu recovery from the Ni-Cu enriched matte by investigating the mechanisms of leaching, as well as identifying the effect of processing variables; namely, temperature and concentration of lixiviant under oxidative and non-oxidative conditions.

The sulphide chemistry is very complex due to the fact that sulphide concentrates usually consist of highly intergrown sulphide minerals. This is one of the reasons why there is a lack of understanding and knowledge of the mechanism of leaching. These processes are thus often not operated at optimum conditions (Rademan, 1999).

Furthermore, the morphology and mineralogy of complex sulphide ores differs in nature, thus there is a need for a separate study for each type of matte. Leaching of Ni-Cu mattes is normally conducted in two stages, with the first stage aimed at removing most of the nickel and cobalt and precipitating the Cu by cementation. The second stage is normally aimed at removal of Cu and the remaining base metals to produce a platinum-rich residue (Lamya and Lorenzen, 2009). Although the entire leaching process can be achieved in a single stage, multiple stage leaching is highly selective, enabling the recovery of the base metals in subsequent processes (Hofirek and Kerfoot, 1992). In addition, higher metal extraction levels are achieved. Twostage leaching studies were therefore conducted in this work.

Methodology The parameters investigated were temperature and concentration of lixiviant. Single- and two-stage leaching processes under oxidative and non-oxidative conditions were conducted. The effect of temperature was investigated using single-stage non-oxidative leaching at temperatures of 50, 70, and 90°C ± 5°C for each experimental run, with a constant Fe3+ concentration of 5 g/L. Non-oxidative twostage leaching was used to investigate the effect of ferric chloride concentration (lixiviant), with varying concentrations of 5, 8, and 11 g/L Fe3+ and a constant temperature of 90°C for each experimental run. The ferric chloride concentration and temperature were the same for both the first- and second -stage leaching. Based on the results of the temperature investigations, a temperature of 90 ± 5°C was chosen for further investigations.

Apparatus and reagents The non-oxidative atmospheric leaching experiments were performed in a 5 L glass vessel fitted with four baffles, a variable-speed overhead stirrer with a flat blade turbine-type impeller, and a heating element with a temperature controller.

The cover of the vessel had two ports for holding a pH/temperature probe and for taking samples. These ports allowed for minimal air ingression into the reactor during the experiment.

The matte was obtained from Impala Platinum Refinery, and ferric chloride was used as the leaching agent.

Hydrochloric acid was used to control the pH below 3.

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Characterization The matte before and after leaching was characterized by Xray diffractometry (XRD) with a Cu Kα radiation source (Xpert Phillips) to determine the mineralogical phases present in the matte, and X-ray fluorescence (XRF) (PW 2540 VRC Sample changer) to determine the elemental composition of the matte. Scanning electron microscopy (SEM) (JOEL JSM 5600) coupled with energy dispersive spectroscopy (EDS) was used to confirm the mineralogical phases in the matte and capture the particle morphology. A laser diffraction technique (Malvern Mastersizer 2000) coupled to a liquid dispersing unit (Hydro 2000G) was used to determine the particle size distribution (PSD) of the matte before and after leaching. Ni and Cu concentrations in the leach liquor were measured by inductively coupled plasma-optical emission spectrometry (ICP-OES (Spectro Arcos Fsh12) and atomic absorption spectroscopy (AAS) (Thermo Scientific ICE 3000 instrument).

Experimental procedure Leaching solution (4 L) was heated in the reactor with agitation until the required temperature was reached; 40 g of matte of particle size -105+75 μm was then added into the vessel. The stirrer was set to the required constant speed of 600 r/min and timing of the experiment started. HCl solution (1.5 M) was added stepwise to the mixture to keep the pH below 3. Temperature and pH at t=0 were recorded and thereafter monitored every 15 minutes. Liquor samples of 25 mL were taken after every 15 minutes for the 200-minute duration of leaching; which was considered a sufficient time to yield a high percentage extraction. The liquor samples were filtered and the filtrate was kept in sample bottles for chemical analysis. At the end of each experiment, the pulp was allowed to cool to room temperature and then filtered.

The filter cake was washed with demineralized water and dried overnight. The residue samples were also kept for chemical analysis.

For non-oxidative two-stage leaching, the general procedure as outlined above was followed. However, after 200 minutes of first-stage leaching, the leaching solution was decanted from the reactor and then filtered. Thereafter 4 L of fresh leaching solution of the desired concentration was then added to the reactor and the filtered cake returned to the reactor to commence second-stage leaching. The concentrations of Fe3+ used in the non-oxidative two-stage leaching tests were 5, 8, and 11 g/L Fe3+. The concentration was kept constant for both first- and second-stage leaching.

For the two-stage oxidative leaching the experimental procedure was similar to that of the non-oxidative two-stage leaching, with the exception of the introduction of oxygen, and a fixed Fe3+ concentration of 11 g/L Fe3+ and a constant temperature at 90 ± 5°C were used based on the findings of the previous experiments. In the first stage, fresh matte was contacted with fresh lixiviant for 200 minutes, and in the second stage, solid residue from the first stage was contacted with fresh solution and further leached for 200 minutes.

Oxygen below 5 kPa was spurged into the solution throughout the 200-minute experimental run. The oxygen pressure was kept constant.

The Journal of The Southern African Institute of Mining and Metallurgy

Atmospheric oxidative and non-oxidative leaching of Ni-Cu matte Results and discussion Matte characterization before leaching reaction The XRD pattern of the matte before leaching is shown in Figure 1. The major phases detected in the matte were djurleite (Cu1.96S) and hazelwoodite (Ni3S2). Hazelwoodite had the highest peak intensity of 100% at 31.0532 angle position, and djurleite had the highest relative intensity of 25% at 32.5175 angle position.

The phases identified by SEM-EDX were Ni alloy, hazelwoodite, djurleite, and some PGMs (Figure 2).

The elemental composition of the matte before leaching as determined by XRF is shown in Table I. Copper constituted 31 mass % while nickel (i.e. Ni alloy and Ni3S2) was the major element in the matte, representing nearly half of the matte. Since the matte is a sulphide, sulphur also constituted a significant amount (13 mass %). The remaining 2% consisted of Fe and Co, while PGMs like Pt and Pd accounted for only 0.5%.

SEM of the matte showed that the particles were irregularly shaped with a smooth outer surface layer i.e. no cracks and veins on the particles (Figure 3).

The PSD of the matte generated by a laser diffraction technique is shown in (Figure 4). The modal size of the matte was 138 μm before leaching. The volume distribution

(Figure 4a) was transformed into the number distribution (Figure 4b) and shows that the matte consisted of a larger number of smaller sized particles, which however; contributed a small volume percentage.

The d(0.1) showed that 10% of the particles were smaller than 86.27 μm before leaching, whereas the d(0.9) demonstrated that 90% of the particles were smaller than 209.58 μm.

Table I

Elemental composition of the matte before leaching, mass % Co

Cu

Fe

Ni

Pb

Pd

Pt

S

Se

Si

Other

0.4

31

1.4

50

0.081

0.21

0.27

13

0.09

1.2

2.35

a = Ni3S2 b = Cu2S

0

10

20

30

40

50

60

70

2θ

80

Figure 3—SEM images of the matte particle before leaching (magnification 100×)

Figure 1—XRD spectra of the matte before leaching

Ni Alloy

Hazelwoodite

Chalcocite

PGMʼs

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Figure 4a—Particle size distribution (by volume) of the matte before leaching VOLUME 114

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Figure 2—SEM-EDX micrographs of the matte before leaching

Atmospheric oxidative and non-oxidative leaching of Ni-Cu matte that almost all the Ni alloy was extracted. This finding agrees with that of Rademan et al. (1999), who stated that Ni alloy was the most reactive phase, followed by Ni3S2, and lastly Cu2S. Micro-cracks on the particles can be seen in Figure 6.

Micro-cracking enhances the leaching rate by increasing the surface area available for reaction; thus increasing the rate of diffusion of the lixiviant and products. The metathesis process was also believed to be taking place simultaneously with cementation according to Equation [1].

Ni3S2 + 2Cu2+ → Cu2S + NiS + 2Ni2+

[1]

Figure 4b—Particle size distribution (by number) of the matte before leaching

Effect of leaching variables Temperature Experiments were conducted at temperatures of 50, 70, and 90°C ± 5°C, at a constant concentration of 5 g/L Fe3+. The recoveries of Ni and Cu as a function of leaching time are shown in Figure 5. Maximum percentage extractions of Ni for temperatures of 50, 70, and 90°C were 18, 27, and 55% respectively. The highest Ni recovery was obtained at 90°C. A temperature of 90°C was therefore chosen as the optimum temperature in this study, since it is advantageous to leach at a temperature near the boiling point of ferric chloride of approximately 105°C (Dutrizac, 1992).

Researchers such as Park et al. (2006), Rademan et al.

(1999) and Dutrizac. (1992) have demonstrated that temperature has a direct influence on leaching. The higher the temperature, the higher the leaching rate, therefore the higher the metal recovery. The results obtained in this study are in agreement with this conclusion. The extraction of copper, however, decreased with increasing temperature (Figure 5b) due to cementation. At 50, 70, and 90°C the maximum percentage extractions of Cu were 14, 4, and 2% respectively. The cementation process was favoured by higher temperatures. Cu extraction was highest at 50°C because the Ni extraction rate was the lowest at this temperature, thus allowing a small amount of copper to be leached. Cu was rejected at this temperature from 14% to approximately 4% in the leach liquor by the end of 200 minutes.

The elemental composition of the matte before and after leaching at 90ºC is shown in Table II. The copper content was 31% before leaching and 42% after leaching. This increase was due to the increased surface exposure of copper sulphide in the residue as a result of leaching. The relative proportion of Ni in the matte decreased by 14% after leaching (Table II). This seemed very low at this stage, but was attributed to the fact that Ni alloy was the phase that leached the most. This finding was substantiated by the SEM-EDS analysis after leaching at 90°C. Figure 6 shows that hazelwoodite and copper sulphide are the remaining minerals in the matte after leaching, thus proving

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Figure 5a—Percentage extraction of Ni after single-stage non-oxidative leaching with 5 g/L [Fe3+]

Figure 5b—Percentage extraction of Cu after single-stage nonoxidative leaching with 5 g/L [Fe3+]

Table II

Quantitative elemental composition of the matte before and after single-stage non-oxidative leaching with 5 g/L [Fe3+] at 90º C, mass%

Before After

Co

Cu

Fe

Ni

Pb

0.4 0.06

31 42

1.4 0.67

50 36.0

0.081 0.213

Pd

Pt

0.21 0.27 0.306 0.405

S

Other

13 7.6

2.35 12.16

The Journal of The Southern African Institute of Mining and Metallurgy

Atmospheric oxidative and non-oxidative leaching of Ni-Cu matte PGMʼs

Hazelwoodite

Chalcocite

Micro cracks

Figure 6—SEM-EDS image of the matte after single-stage non-oxidative leaching with 5 g/L [Fe3+] at 90°C

mineralogical phases containing Ni after leaching with Fe3+ 11 g/L as NiS, Ni3S2, and Ni3S4 (Figure 9).

Ni7S6 was, however, detected only after leaching with 5 g/L Fe3+. This can be explained by the fact that this Fe3+ concentration resulted in the slowest leaching rate, allowing the quasi-intermediate product Ni7S6 (a mineral species that is stable for only a short period during the reaction) to be detected. XRD also detected elemental sulphur; this is the sulphur produced from the possible oxidation of NiS (Equation [2]).

According to Rademan et al. (1999) base metals are gradually leached out of the sulphide lattice to form species with lower metal-to-sulphur ratios. It is proposed that the transformation proceeded as follows: Ni3S2 is rapidly altered to Ni7S6, then forms NiS, and finally Ni3S4. Sulphide minerals

However, the XRD analysis (Figure 7) shows that no mineral transformation took place during the leaching process.

The inability of XRD to detect NiS can be attributed to three reasons: (1) the leaching reaction had not run to completion, (2) the NiS concentration was too low to be detected, or (3) NiS was oxidized by ferric ions according to Equation [2] thus liberating elemental sulphur.

NiS + 2Fe3+ → Ni2+ + 2Fe2+ + S0

a = Ni3S2 b = Cu2S

[2]

Table II also shows that other base metals, such as Co and Fe, are oxidized by Fe3+, depleting the concentration of Fe3+. These elements show a decrease in concentration after leaching (Table II). Little attention was paid to these metals in this study because they are present in small amounts compared to Ni and Cu, and their mineralogical phases were not detectable by either SEM or XRD. Based on the initial findings, it was proposed that to obtain maximum Ni and Cu extraction, it is advantageous to leach in two stages at 90°C. The first stage results in a higher extraction of Ni. Secondstage leaching was carried out using fresh leaching solution on the partially leached solid residue to extract the remaining Ni and Cu. It was also concluded that the ferric ion solution should be prepared at higher concentrations, so that the Ni and Cu dissolution rates and extractions are not compromised by other Fe3+-depleting agents.

0

20

40

60

80

2θ

Figure 7—XRD pattern after single-stage non-oxidative leaching with 5 g/L [Fe3+] at 90°C

Concentration

The Journal of The Southern African Institute of Mining and Metallurgy

Figure 8—Percentage extraction of Ni after two-stage non-oxidative leaching with varying [Fe3+] at 90°C

Table III

Quantitative elemental analysis after two-stage nonoxidative leaching with 11 g/L [Fe3+] at 90ºC, mass% Co Before After

Cu

Fe

Ni

Pb

Pd

Pt

S

Other

0.4 31 1.4 50 0.081 0.21 0.27 13 0.074 46.63 7.54 17.38 0.104 0.184 0.231 5.84

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The concentrations of Fe3+ investigated were 5, 8, and 11 g/L, at a constant temperature of 90°C. It was therefore advantageous to leach in two continuous stages because the percentage Ni extraction increased from 18% for single-stage leaching (Figure 5) to 26% (Figure 8) for two-stage leaching at 90°C using 5 g/L Fe3+.

The Ni percentage extractions for 5, 8, and 11 g/L of Fe3+ were approximately 26, 29, and 62% respectively. Thus further tests to establish the leaching mechanisms and particulate processes were conducted for samples obtained when leaching with 11 g/L Fe3+.

XRF analysis (Table III) showed a decrease in initial Ni content of the matte from 50% to approximately 17% after leaching.

Sulphur mass percentage also decreased, implying that sulphur was liberated into the solution. XRD identified the

Atmospheric oxidative and non-oxidative leaching of Ni-Cu matte can be solubilized in acid or basic medium but form elemental sulphur only under oxidizing conditions. The leaching of sulphides in acid medium in the presence of ferric chloride takes place readily with the liberation of elemental sulphur (Havlík, and Kammel, 1995). The tests performed in this study were non-oxidative; but elemental sulphur was liberated, as shown in Figure 9. This demonstrates that ferric ions performed two functions – one of oxidizing the sulphide lattice and the other of attacking (acid leaching) the ore. This finding agrees with that of Park et al. (2006). The only disadvantage with non-oxidative ferric chloride leaching (Equation [3]) is that the ferrous ions formed cannot be regenerated back to ferric ions according to Equation [4]:

MS + 2Fe3+ → M2+ + 2Fe2+ + S M = Cu, Ni 2Fe2+ + 2H+ + 0.5O2 → 2Fe3+ + H2O

[3] [4]

The leaching of copper sulphides is slightly more complex than that of nickel sulphide, as seen in the single-stage temperature leaching studies. A similar pattern is observed in the two-stage leaching process, where higher extractions are obtained in the first few minutes of the experiment, followed

by Cu cementation as early as 40 minutes into the reaction time (Figure 10). At the end of the two-stage leaching

a = Ni3S4 b = NiS c = Cu1.8S d=S e = Ni3S2 f = Cu g = Cu2S

process using 5, 8, and 11 g/L of Fe3+, the Cu extractions were 3, 11, and 11% respectively. However, the highest extractions, of 35, 67, and 96% for Fe3+ concentrations of 5, 8, and 11 g/L respectively, were obtained in the first 15 minutes, before the cementation process began.

Table III shows an increase in the Cu content; this is because a Cu–rich cake was created as a result of cementation, while more Ni was dissolved into the solution.

Rademan et al. (1999) stated that Cu2S is leached to form digenite (Cu1.8S), with (Cu31S16) forming as a quasiintermediate product. The digenite leaches further to form covellite (CuS). Figure 9 shows that Cu2S was transformed to Cu1.8S, with Cu31S16 detected only with leaching at 5 g/L Fe3+. The leaching of copper proceeded as shown in Equations [5–7]:

5Cu2S + 2Fe3+ → Cu1.8S + Cu2+ + 2Fe2+ Cu1.8S + 8Fe3+ → 5CuS + 4Cu2+ + 8Fe2+ CuS + 2Fe3+ → Cu2+ + 2Fe2+ + S

[5] [6]

[7]

However, CuS formed by Equation [6] was not detected

by XRD (Figure 9), possibly as a result of copper rejection and the slow leaching rate. An oxidative two-stage leach was therefore conducted at the selected temperature of 90°C and 11 g/L Fe3+ concentration to increase the overall leaching rate and base metal extraction. During first-stage leaching i.e. the leaching of nickel alloy, micro-cracking of the particles occurred, exposing the Ni3S2 and Cu2S on the edges of the cracks (barely visible) and on the outside of the particles (Figure 6), thus increasing the surface area available for leaching. In the second stage the size of the micro-cracks increased as the leaching process continued until the particles becomes lightly porous, depositing PGMs on their surface, and finally particle breakage occurred as shown in Figure 11.

Effect of oxygen 0

20

40

60

80

2θ

Figure 9—XRD pattern after two-stage non-oxidative leaching with 11 g/L [Fe3+] at 90°C

According to Qui et al. (2007) oxygen diffuses initially into the solution from the gas/liquid interface and then diffuses further into the solid/liquid interface. Oxygen participates in the reaction after it contacts the ore surface. The solubility of oxygen in water is affected by the temperature and partial pressure. Increasing the partial pressure of oxygen increases its solubility. However; Deng et al. (2001) stated that

Copper Sulphide

Hazelwoodite

Porous particle

PGMʼs

Micro cracks

Figure 10—Percentage extraction of Cu after two-stage non-oxidative leaching with varying [Fe3+] at 90°C

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Figure 11—Electron micrograph of the matte after two-stage nonoxidative leaching with 11 g/L [Fe3+] at 90°C The Journal of The Southern African Institute of Mining and Metallurgy

Atmospheric oxidative and non-oxidative leaching of Ni-Cu matte leaching at a temperature above 85°C is not desirable, because oxygen solubility in the slurry decreases at higher temperatures, especially near the boiling point of the solution. Since the oxygen partial pressure for this study was low and the leaching temperature was high, oxygen solubility was enhanced by the high impeller speed of 600 r/min, leading to good dispersion and dissolution of oxygen in the leach slurry. Figure 12 illustrates how effective and important oxygen is to the leaching process.

Approximately 98% Ni was extracted under oxidative leaching. Ni alloy was extracted simultaneously with almost all of the Ni3S2. Under non-oxidative conditions, only a total of 62% Ni was extracted. Nickel was dissolved via metathesis (Equation [9]) and oxidation (Equations [10–12]).

2Fe3+ + 3H2O → Fe2O3 + 6H+ Ni3S2 + 2Cu2+ → Cu2S + NiS + 2Ni2+ Ni0 + 2H+ + 0.5O2 → Ni2+ + H2O 3Ni3S2 + 4H+ + O2 → Ni7S6 + 2Ni2+ + H2O Ni7S6 + 2H+ + 0.5O2 → 6NiS + Ni2+ + H2O

[8] [9] [10] [11] [12]

Equation [10] shows that oxygen is responsible for Ni alloy dissolution, and copper rejection is a result of Ni metathesis (Equation [9]). The intermediate product Ni7S6 formed (Equation [11]) quickly transformed into NiS (Equation [12]). This confirms the findings of Rademan et al.

(1999). It can thus be concluded that oxygen increases liberation from the intricate sulphide bonds, thus enhancing the dissolution process. Cu extraction from a copper–nickel complex sulphide ore was more complex than Ni extraction.

Figure 13 depicts the extraction of Cu. The leaching chemistry changed and the solubilized Cu was re-deposited on the residue. It was evident that under non-oxidizing conditions Cu is extracted rapidly and suddenly undergoes cementation. The same process occurred under oxidizing conditions. The nickel remaining in the partially leached residue of the first stage was further dissolved by metathesis and oxidation in the second-stage leaching. It was concluded that no copper was leached from matte in the two stages. The majority of the copper originally present in the matte remains intact through the oxidative leaching, because copper is not as easily oxidized as nickel and as a result of copper cementation by nickel. To enhance the extraction of Cu, a third oxidative leaching stage is necessary, as Figure 5 shows that in the absence of cementation most of the Cu is extracted within the first 50 minutes of leaching. Chalcocite will be leached and

transformed to covellite (CuS). It is predicted that copper sulphide would be attacked by ferric ions (Equations [13–16]), and oxygen in this instance would serve as an oxidant responsible for the regeneration of ferrous ions to ferric ions, according to Equation [13].

2Fe2+ + 2H+ +0.5O2 → 2Fe3+ + H2O 5Cu2S + 2Fe3+ → Cu1.8S + Cu2+ + 2Fe2+ Cu1.8S + 8Fe3+ → 5CuS + 4Cu2+ + 8Fe2+ CuS + 2Fe3+ → Cu2+ + 2Fe2+ + S

[13] [14] [15] [16]

After the first-stage non-oxidative leaching, microcracking of the matte was observed (Figure 6), and after the two-stage non-oxidative leaching, extensive microcracking occurred with minor pores formed on the particles (Figure 11). Figure 14 shows extensive pores formed on the particle under two-stage oxidative leaching, exposing more PGMs. It is believed that in the third stage, the particles would proceed to complete breakage as a result of increased oxidation and dissolution.

The XRF results agreed with those from SEM, as they showed that after oxidative leaching the relative proportions of Pd and Pt in the matte was increased enabling them to be detected by XRF (Table IV).

Figure 13—Cu extraction at 90°C with 11 g/L [Fe3+]

Chalcocite

Hazelwoodite

Porous particle

PGMʼs

The Journal of The Southern African Institute of Mining and Metallurgy

Figure 14—Electron micrograph of the matte after two-stage oxidative leaching with 11 g/L [Fe3+] at 90°C VOLUME 114

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Figure 12—Ni extraction after two-stage oxidative and non-oxidative leaching with 11 g/L [Fe3+] at 90°C

Atmospheric oxidative and non-oxidative leaching of Ni-Cu matte Table IV

Quantitative elemental analysis after two-stage, oxidative leaching with 11 g/L [Fe3+] at 90ºC, mass% Co

Cu

Fe

Ni

Pb

Pd

Pt

S

Other

Before 0.4 31 1.4 50 0.081 0.21 0.27 13 2.35 After oxidation 0.068 48.0 1.90 26.0 0.090 0.350 0.560 6.90 16.13

[1–(1–X)1/3]2 = k1t [1–3(1–X)2/3 + (1–X) + α (1–(1–X)1/3)]2 = k2t

[17] [18]

where X is the fraction reacted, k1 and k2 are reaction rate constants, and α is a constant.

These results show that the leaching mechanism is diffusion controlled as result of the formation of a porous sulphur layer on the surface of the particles during leaching.

4000

Since XRF expresses the elemental compositions as fractions of the total mass of the sample, during leaching some constituents of the matte are removed, changing the relative proportion of the elements in the matte Copper mass% increased as a result of cementation, while Ni content decreased because of dissolution. Sulphur was also liberated into the solution; however, no evidence of sulphur passivation in the leaching process, as observed by Park et al. (2006), was found in this study. XRD did not detect any hazelwoodite (Ni3S2) or chalcocite (Cu2S) in the leached residue (Figure 15). The Ni transformed as far as NiS and a part of the Ni was oxidized to NiO. XRD confirmed that copper underwent dissolution in the first 45 minutes, as the copper phase detected was Cu1.8S.

a = NiS b = Cu1.8S c = NiO

3500 3000 2500 2000 1500 1000 500 0 0

10

20

30

40

50

60

70

80

2θ

Figure 15—XRD pattern after two-stage oxidative leaching with 11 g/L [Fe3+] 90°C

Particulate processes The PSD of the matte obtained by a laser diffraction technique is shown in Figure 16.

The peaks represent the modal size (the particle size with the highest volume percentage). The modal size of Ni-Cu matte before leaching was 138 μm. The modal size after leaching at 90°C using 5 g/L Fe3+ (single-stage LX) reduced to 107 μm. This suggests a decrease in particle size either as a result of dissolution or breakage of the particles in the larger size fraction. There was no significant shift in the volume distribution and modal size after two-stage nonoxidative leaching using 11 g/L Fe3+; however, there was a decrease in the volume percentage of the mode. This indicates a decrease in the proportion of larger particles as the number of leaching stages was increased. The modal size after oxidative leaching was the same as that reported after nonoxidative leaching, but shifted slightly to the left and decreased in height, implying a greater number of particles decreased in size.

Figure 16—PSD (by volume) curves for the matte before and after leaching

Kinetics Various researchers such as Fan et al. (2010) and Jin et al. (2009) have found that the dissolution kinetics of NiCu matte during the leaching process follow the shrinking core model (SCM). Only nickel dissolution kinetics were investigated in this study, because nickel was the only metal dissolved during the leaching process since copper was precipitated. The shrinking core model gave a satisfactory fit to the experimental data only at 90°C, and the most suitable model was the product layer/ash diffusion model (Equation [17]) which had a correlation coefficient of 0.8 (Figure 17a).

The experimental data at 90°C was fitted to the threedimensional diffusion model by Jander (Equation [18]) and yielded a correlation coefficient of 0.93 (Figure 17b):

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Figure 17a—Plot of the mixed (surface and product layer/ash diffusion) controlled process at various temperatures The Journal of The Southern African Institute of Mining and Metallurgy

Atmospheric oxidative and non-oxidative leaching of Ni-Cu matte

Figure 17b—Plot of the data at 90°C on the three-dimensional diffusion model by Jander

Conclusion Ni was selective leached from a Ni-Cu matte using a twostage oxidative leaching process at 90°C and 11 g/L Fe3+. The Ni extraction was 98%. Non-oxidative leaching was found to be less effective than oxidative leaching, as only 62% of the Ni was extracted under the same conditions. Cu was entirely rejected from solution by cementation, which would enable it to be selectively extracted in a third leaching stage. Oxygen was found to enhance the liberation of metals from the intricate sulphide bonds, thus enhancing the dissolution process. The major disadvantage of non-oxidative leaching is that ferrous ions can not be regenerated back to ferric ions, leading to lower metal extractions. The shrinking core model did not give an acceptable fit to the experimental data at temperatures below 90°C, with the three-dimensional diffusion model by Jander giving the best fit at 90°C. Leaching resulted in extensive micro-cracking and pore formation on the particle surfaces, resulting in particle breakage. Although a porous sulphur layer was formed on the surface of the particles during leaching, sulphur passivation was not found to occur.

Acknowledgements The authors would like to thank Impala Platinum for supplying the Ni-Cu matte, and the University of Johannesburg URC for financial support.

References AGRAWAL, A., KUMAR, V., PANDEY, B.D., and SAHU, K.K. 2006. A comprehensive review on the hydro metallurgical process for the production of nickel and copper powders by hydrogen reduction. Materials Research Bulletin, vol. 41. pp. 879–892.

BRUGMAN, C.F. and KERFOOT D.G. Treatment of nickel-copper matte at Western Platinum by the Sheritt acid leach process. Proceedings of the 25th Annual Conference of Metallurgists, Toronto, Canada, 17-20 August 1986.

Canadian Institute of Mining, Metallurgy and Petroleum, Toronto. pp. 512–531.

The Journal of The Southern African Institute of Mining and Metallurgy

DENG, T., LU, Y., WEN, Z., and LIU, D. 2001. Oxygenated chloride-assisted leaching of copper residue. Hydrometallurgy, vol. 62, no. 1. pp. 23–30.

DUTRIZAC, J.E. 1992. The leaching of sulphide minerals in chloride media.

Hydrometallurgy, vol. 29, no. 1–3. pp.1–45.

FAN, C., LI, B., FU, Y., and ZHAI, X. 2010. Kinetics of acid-oxygen leaching of low-sulfur Ni-Cu matte at atmospheric pressure. Transactions of Nonferrous Metal Society of China, vol. 20, no. 6. pp. 1166–1170.

FORWARD, F.A., and MACKIW, V.N. 1955. Chemistry of the ammonia pressure process for leaching Ni, Cu, and Co from Sheritt Gordon sulphide concentrates. Journal of Metals, vol. 7. pp. 457–463.

HABASHI, F. 1993. A Textbook of Hydrometallurgy. Metallurgie Extractive Quebec. Quebec, Canada. p. 99.

HAVLÍK, T. and KAMMEL, R. 1995. Leaching of chalcopyrite with acidified ferric chloride and carbon tetrachloride addition. Minerals Engineering, vol. 8, no. 10, pp. 1125–1134.

HOFIREK, Z., and KERFOOT, D.G.E. 1992. The chemistry of the nickel-copper matte leach and its application to process control and optimization.

Hydrometallurgy, Theory and Practice, Proceedings of the Ernest Peters International Symposium. Cooper, W.C. and Dreisinger, D.B. (eds.).

Hydrometallugy, vol. 29. pp. 357–381.

JIN, B., YANG, X., and SHEN, Q. 2009. Kinetics of copper dissolution during pressure oxidative leaching of lead-containing copper matte.

Hydrometallurgy, vol. 99, no. 1–2. pp.119–123.

LAMYA, R.M., and LORENZEN, L. 2009. A semi-empirical kinetic model for the atmospheric leaching of a Ni-Cu converter matte in copper sulphatesulphuric acid solution. Journal of the Southern African Institute of Mining and Metallurgy, vol. 109. pp. 755–760.

PARK, K.H., MOHAPATRA, D., and REDDY B.R. 2006. A study on the acidified ferric chloride leaching of a complex (Cu–Ni–Co–Fe) matte. Separation and Purification Technology, vol. 51, no. 3. pp. 332–337.

PEARCE, R.F., WARNER, J.P., and MACKIW, V.N. 1960. A new method of matte refining by pressure leaching and hydrogen reduction. Journal of Metals, vol. 12. pp. 28–31.

PLASKET, R.P. and ROMANCHUK, S. 1978. Recovery of nickel and copper from high grade matte at Impala Platinum by the Sherritt Process.

Hydrometallugy, vol. 3. pp. 135–151.

QUI, T., NIE, G., WANG, J., and CUI, L. 2007. Kinetic process of oxidative leaching of chalcopyrite under low oxygen pressure and low temperature.

Transactions of Nonferrous Metal Society of China, vol. 17, no. 2. pp. 418–422.

RADEMAN, J.A.M., LORENZEN, L., and VAN DEVENTER, J.S.J. 1999. The leaching characteristics of Ni–Cu matte in the acid–oxygen pressure leach process at Impala Platinum. Hydrometallurgy, vol. 52, no. 3. pp. 231–252. ◆ VOLUME 114

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For a better fit it is recommended that the PSD of the matte be included in the SCM. Due to non-fitting models at lower temperatures, the activation energy could not be determined.

For further information contact: Raymond van der Berg, Head of Conferencing SAIMM · P O Box 61127, Marshalltown 2107 Tel: +27 11 834-1273/7 · Fax: +27 11 833-8156 E-mail: raymond@saimm.co.za Website: http://www.saimm.co.za

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Universities and decision-making: programme and qualification mix — four learning pathways

by W.P. Nel*

The introduction of the Higher Education Qualifications Framework (HEQF) and the updated Higher Education Qualifications Sub-Framework (HEQSF) has caused many South African university departments to rethink their programme qualification mixes (PQMs). In addition to the requirements stated in the HEQSF, a number of other factors have to be taken into consideration by a university department. These factors include, for example, the standards generated by the Engineering Standards Generating Body (ESGB) and subsequently approved by the Engineering Council of SA (ECSA) and the need to prepare students for various categories of professional registration with ECSA. This means that a university department has to choose the correct mix of Learning Programmes (LPs) from the HEQSF menu (which consists of 13 types of LPs). Preparing students for ECSA registration is aligned with the mission of universities, which is to teach and undertake research. However, research and the LPs associated with research go beyond the requirements for current ECSA registration. Assuming that universities offering engineering LPs would elect to prepare students for both ECSA registration and teach them to produce research outputs, which is mostly done at Master and Doctorate levels (NQF Levels 9 and 10), then it follows that academics are more interested in NQF Level 5 to 10 pathways (abbreviated as ‘L5-10’) rather than the shorter pathways required towards professional registration. (For example, ECSA requires an NQF L5-L7 pathway for registration as a candidate professional technologist. This specific pathway may consist, for example, of two LPs, namely the 360-credit Diploma and the Advanced Diploma.) A L5-L10 pathway is a combination of LPs that will prepare the learner with a NSC (or equivalent qualification at level 4) to Doctoral level (level 10). Universities may choose at least four major pathways from the HEQSF menu in order to educate and develop students from NQF Level 5 to 10. However, various pathways towards registration in the category of candidate with ECSA are also embedded into these four NQF L5-L10 pathways, where each consist of a unique combination of LPs. Each of these pathways has an opportunity cost, and economic reality means that smaller departments may have to choose between the four pathways. Of all the many factors involved in PQM decision-making, the focus of this paper is on the HEQSF requirements, ECSA standards, and ECSA registration and how these, together with the various qualifications and educational LPs provided for by the HEQSF may impact on the PQM decision taken by engineering departments and schools at South African universities. The proposed four NQF L5-L10 ‘pathway tool’ for PQM decision-making may be useful for pointing out the advantages, disadvantages, and applications of the various pathways and combinations of pathways. Rather than deciding from a menu of thirteen qualifications and associated LPs, this article proposes that decision-making be undertaken on the basis of a menu of four main articulated ‘NQF L5-L10’ pathways (which also include one or more of the ECSA’s pathways for professional registration). The proposed ‘NQF L5-L10 pathway’ tool is an attempt to move one step closer to the aim of achieving a structured decision-making approach for designing a PQM at departmental level.

The Journal of The Southern African Institute of Mining and Metallurgy

The ECSA designed standards for some of the 13 LPs that form part of the HEQSF ECSA have to date not developed competency standards for Levels 9 (Masters) and L10 (Doctorate). If universities design LPs according to these standards, then learners would be eligible to comply with ECSA’s educational requirements to register in the categories of candidate Pr Techn., Pr Tech., Pr Cert. Eng., and Pr Eng. Keywords Higher Education Qualifications SubFramework (HEQSF), educational learning programmes (LPs), educational pathways, programme and qualification mix (PQM), registration in the appropriate candidate category with the ECSA approved standards, PQM decision-making, articulation.

Introduction Many university departments are rethinking their programme and qualification mixes (PQMs) as a result of the implementation of the new Higher Education Qualifications SubFramework (HEQSF). Deciding on a viable PQM for a university department is a complex matter that is influenced by many different factors, requirements and constraints. In addition to the requirements stated in the HEQSF, many other factors, such as market requirements, economic viability, path dependency, articulation, mode of teaching (online learning vs classroom-based education), and university type (e.g. compre-

* University of South Africa (UNISA). © The Southern African Institute of Mining and Metallurgy, 2014. ISSN 2225-6253. Paper received Feb. 2014; revised paper received Apr. 2014.

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Synopsis

Universities and decision-making: programme and qualification hensive, research-intensive, or technology-focused) also have to be considered. One of the implications of the mergers of Rand Afrikaans University with Technikon Witwatersrand to form the University of Johannesburg and the old University of South Africa (Unisa) with Technikon Southern Africa to form the new Unisa is that these comprehensive universities may now offer the B Eng degree. Should they do that? This decision forms part of the decision-making that will determine their future PQMs, the combination of HEQSFaligned programmes to offer in the future.

Competency standards approved by ECSA are the prime requirement to be considered by departments offering engineering-related programmes and an important factor when deciding on a PQM. The inclusion or exclusion of Work-Integrated Learning (WIL) is also an important HEQSFrelated provision that has to be taken into consideration by university departments.

The word ‘pathway’ is to be found in a number of ECSA documents. ‘Pathway’ means the combination of educational LPs that must be followed to attain a specific goal (e.g. compliance with the educational requirements for professional registration as engineer, certificated engineer, technologist, and technician). Academic staff at universities have two major responsibilities: tuition and research.

Research and the LPs associated with research go beyond the requirements for current ECSA registration. Assuming that universities offering engineering LPs would elect to both prepare students for ECSA registration and teach them to produce research outputs, which is mostly done at Masters and Doctorate Levels (NQF Levels 9 and 10), then it follows that academics are more interested in NQF Level 5 to 10 pathways (abbreviated as ‘L5-10’) rather than the shorter pathways required towards professional registration. (For example, ECSA requires an NQF L5-L7 pathway for registration as a candidate professional technologist. This specific pathway may consist, for example, of two LPs, namely the 360-credit Diploma and the Advanced Diploma.)

A L5-L10 pathway is a combination of LPs that will prepare the learner with a National Senior Certificate (NSC) (or equivalent qualification at exit level 4) to Doctoral level (exit level 10).

This paper does not focus on all the factors, requirements, and constraints that may impact, in various ways, on the PQM decision, but attempts to investigate the proposed ‘NQF L5-L10 pathways method’ or tool to simplify PQM decisionmaking while considering only some of the factors, requirements, and constraints such as the requirements specified by the ECSA and the HESQF. For this purpose, I have identified four main pathways (for NQF L5-L10) from the HEQSF document for developing and educating someone from National Senior Certificate (NSC) to doctoral level. In this article, I shall show that these NQF Level 5 to 10 pathways, combined with the various pathways leading to registration with the appropriate candidate category with the ECSA, are important tools that can be used in the complex PQM decision-making process. These pathways are proposed as a method for reducing the complexity of decision-making (from the broad menu of thirteen HEQSF qualifications) to

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one of deciding between articulated pathways to achieve various goals while considering various factors, requirements, and constraints that may impact on the PQM decision of a department. In essence, I have generalized and expanded ECSA’s pathways for professional registration to a proposed ‘pathway tool’ in order to reduce the complexity of PQM decision-making by one level. Instead of deciding from a menu of qualifications and associated programmes allowed by the HEQSF, I shall propose, in this article, that decisionmaking be done from a menu of articulated NQF L5-L10 pathways that incorporate HEQSF-compliant ECSA pathways for engineering practitioners.

Background information A PQM of a university is a list, menu, or mix of approved LPs and qualifications that will be subsidized by the Department of Higher Education and Training (DHET). A university department or school has to follow a certain procedure to get its programmes on to such a list. In the future, the PQM of South African universities may consist only of programmes provided for by the HEQSF. The HEQF was gazetted by the South African Minister of Education in 2007 and is an integral part of the National Qualifications Framework (NQF). It was updated in January 2013 and is now called the HEQSF. Universities are allowed to offer qualifications at NQF Levels 5 through 10. The thirteen types of qualifications that form part of the HEQSF menu are listed in Table I. It is important to note that Universities of Technology (UoTs) – the former technikons – and comprehensive universities are particularly affected by the HEQSF because national diplomas have been replaced by diplomas and the B Tech-degree has been excluded from the HEQSF. See Tables I and II. ECSA designed standards for some of the 13 LPs that form part of the HEQSF. ECSA has, however, not developed competency standards for the Level 8 Postgraduate diploma, Level 9 Professional Masters, and L10 Professional Doctoral degrees. The HEQSF has significant implications for educational providers. Changes to Work-integrated Learning (WIL) are also described in the HEQSF (2013, p. 11).

A number of South African universities have designed their new HEQSF-aligned PQMs and are in the process of obtaining approval. It may still take a while until the HEQSFaligned PQMs will be implemented. Currently, the requirements for registration with the ECSA for various types of candidacy programmes are as described in Table II. In the case of the technician, certificated engineer, and engineer a single LP is required to meet the academic requirements for professional registration. In the case of the technologist this pathway is longer in terms of the number of LPs that must be completed - it consist of at least two LPs, namely a National Diploma and BTech. See Table II. In the past the UoTs offered cooperative education, meaning that the educational institution and industry cooperated to provide a joint educational programme, which might have included work-integrated learning (WIL). This practice has been largely continued by UoTs and comprehensive universities (CUs) that offer vocational programmes.

The Journal of The Southern African Institute of Mining and Metallurgy

Universities and decision-making: programme and qualification Table I

The HEQSF qualifications menu Undergraduate Qualification

Postgraduate Level

Higher certificate

5

Advanced certificate Diploma (240 credits) Diploma (360 credits)

6 6 6

Advanced diploma Bachelor’s degree (Professional) Bachelor’s degree

Qualification

Level

Postgraduate diploma Bachelor honours degree

8 8

7 7

Master’s degree (Professional) Master’s degree

9 9

8

Doctoral degree (Professional) Doctoral degree

10 10

Table II

Pre-HEQSF pathways towards professional registration Category of professional registration

Pre-HEQSF academic requirements for professional registration in the various categories

Technician Certificated Engineer

National Diploma Recognised certificate of competency (COC) for example the COC Mine Manager (Mine Health and Safety) (Metalliferous)

Technologist

National Diploma + B Tech (Eng), or M Tech (Eng) and pre-requisite LPs

Engineer

B Eng/BSc (Eng)

Source: ECSA (http://www.ecsa.ac.za)

The Journal of The Southern African Institute of Mining and Metallurgy

Four articulated (NQF L5-10) pathways under the HEQSF Figure 1 illustrates four main L5-L10 higher educational pathways, with the National Senior Certificate (NSC) or its equivalent as the admission requirement and the Doctorate as the exit level (Actually, there are more than four such pathways if the differences at Master’s and Doctoral level are also considered.) Please note that Figure 1 relates to engineering, since it uses some of ECSA’s names for the various HEQSF qualifications.

If the sole purpose is to develop a student from NSC to Doctoral level (and not for a specific ECSA registration catergory), then pathway 1 is a substitute for pathways 2, 3, and 4. Similarly, pathway 2 is a substitute for pathways 1, 3, and 4, and so on (for developing a student with a NSC right up to Doctoral level). Note that each qualification in a specific pathway complements the others in the same pathway. This means that all the qualifications, except for the Doctorate, are prerequisites for (subsequent) others in the same pathway.

The removal of any one of the qualifications in a specific pathway will result in articulation deficiency. The view that some programmes or pathways may be substitutes for others points to the fact that they may compete for the same students, assuming that such students meet the admission requirements of the different programmes.

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CUs like UJ and Unisa may in future for example, decide to include both traditional academic and vocation-based programmes in their PQMs. The HEQSF is one of three sub-frameworks. The others are the framework for General and Further Education Subframework (GFETSF) and the framework for Occupational Qualifications (QCTO [sa], p. 3). The HEQSF provides for three progression routes, namely vocational, professional, and general academic. These are listed in Table III and are related to some extent to the ‘types’ of universities found in South Africa. According to the HEQSF (2013, p. 9), undergraduate certificates and diplomas are usually found within the vocational route, while the professional Bachelor, Master’s, and Doctorate degrees are characteristic of the professional route. The general academic learning route focuses primarily on theoretical knowledge and research at higher levels.

It is important to note that the HEQSF specifies certain minimum requirements for qualifications. The minimum number of credits may, however, be exceeded by universities.

For example, the HEQSF specifies that the Higher Certificate should consist of at least 120 credits. ECSA’s minimum requirement is, however, 140 and multiples of 140 credits.

See Table IV (HEQSF, 2013, pp. 21, 22, 25, 28, 30; Van Niekerk, 2013: slide 11). It is not currently clear whether (public) universities will receive subsidy for those credits that exceed the minimum number as specified by the HEQSF.

Universities and decision-making: programme and qualification Table III

Vocational (V), professional (P), and general (G) academic routes provided for in the HEQSF (ECSA qualification standards exist for those in italics) NQF level

Type of qualification and route according to the HEQSF

V

5

Higher Certificate – ‘primarily vocational’ (HEQSF, 2013: 21)

X

LPs approved by the ECSA

6

Advanced Certificate – ‘primarily vocational’ (HEQSF, 2013: 21); ‘particular career or professional context’ (HEQSF, 2013: 23)

X

6

Diploma (minimum, 240 credits) and Diploma (minimum, 360 credits) – ‘primarily has a vocational orientation, which includes professional, vocational, or industry specific knowledge’ (HEQSF, 2013: 24)

X

7

Advanced Diploma – ‘vocational or professional preparation or specialisation’ (HEQSF, 2013: 26)

X

7

(General) Bachelor’s degree – minimum 360 credits (HEQSF, 2013: 26)

8

Professional Bachelor’s degree – minimum 480 credits (HEQSF, 2013: 26)

X

8

Bachelor Honours degree – ‘broad and generic areas of study, disciplines or professions’ (HEQSF, 2013: 30)

X

X

8

Postgraduate Diploma (HEQSF, 2013: 31)

X

X

9

Master’s degree (HEQSF, 2013: 32) (Professional) Master’s degree (HEQSF, 2013: 33)

X

The introduction of the HEQF set into motion a chain of events. The ECSA approved a number of generic HEQF/HEQSF-aligned competency standards for LPs that form part of the pathways for professional registration as an engineer, technologist, technician, certificated engineer, and the category of candidate with the Engineering Council of South Africa (ECSA). The Engineering Standards Generating Body (ESGB) is a committee of ECSA that develops and recommends relevant competency standards (qualification) for engineering practitioners for approval by to ECSA (Van Niekerk, 2013: slide 8).

ECSA pathways obviously have an influence on a faculty or school of engineering’s PQM and the individual departments and/or sections that make up this faculty or school. In a number of cases, the ECSA introduced additional requirements (in addition to those required by the HEQSF) for those qualifications that form pathways to the various categories of professional registration. ECSA’s professional development model towards registration is a two-stage process. Obtaining a relevant, engineering-accredited qualification is the first stage. Stage 2, (professional development of engineering practitioners) consists of a candidacy programme. The various forms of professional registration (and their pathways) with the ECSA are shown in Figure 2 prescribing learning outcomes.

10

Doctoral degree (HEQSF, 2013: 36) (Professional) Doctoral degree (HEQSF, 2013: 38)

P

G

The concept of articulation is an important factor that should guide PQM decision-making. The White Paper for Post-school Education and Training stresses that articulation must be provided between various qualifications, that students should not experience any ‘dead ends’, and that people should be able to improve their qualifications without unnecessary repetition/duplication (DHET, 2013: viii).

X

X

X

X X

Note that the HEQSF provides for two variants of the diploma, one consisting of 240 credits and another consisting of 360 credits (HEQSF, 2013, p. 24).

Figure 1—From National Senior Certificate (NSC) to PhD - four articulated pathways under the HEQSF (ECSA qualification standards exist for those in italics and underlined)

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Universities and decision-making: programme and qualification Table IV

Differences between the HEQSF’s and ECSA’s requirements for selected qualifications Qualification

HEQSF’s minimum specified NQF credits

ECSA’s minimum required NQF credits

Higher Certificate

120

140 – HCert(___Eng)

Advanced Certificate

120

140 – AdvCert(___Eng)

Diploma

240 360

280 - Dip (Eng Tech) 360 – Dip (Eng)

Advanced Diploma

120

140 – AdvDip(Eng)

B degree B degree (Professional)

360 480

420 – BEng Tech 560 – BEng

Bachelor Honours

120

140 – BEng Tech(Hons)

Figure 2—Pathways towards professional registration with the ECSA (Van Niekerk, 2013: slide 12)

‘NQF L5-10’ pathways as a decision-making tool Departments have one of the following three options as far as the selection of NQF L5-10 pathways are concerned: ➤ To offer part of a NQF L5-10 pathway only. In this case, the four main ‘NQF L5-10’ LP are not useful as a tool. In a specific area of study a department may decide to offer one or two programmes on the ECSA The Journal of The Southern African Institute of Mining and Metallurgy

framework (Figure 2) only. This may be done for a number of reasons (e.g. addressing a specific industry need). Another reason may be that such an area of study may not be earmarked for research activities. A decision has been made, for example, at Unisa’s Department of Electrical and Mining Engineering, to offer a Higher Certificate and Advanced Certificate in Mine Surveying only. (Note that PLATO is the professional body for mine surveyors.) ECSA’s competency standards do not cater for programmes in the area of mine surveying, since surveying is not primarily considered as an engineering field of study ➤ To offer one ‘comprehensive’ (NQF L5-10) pathway only. The depth of the field of study may, for example, justify offering one comprehensive pathway. Student numbers, economic viability, staff capacity, and other factors may create an environment in which it is advisable to offer one comprehensive pathway only.

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It is important to note that ECSA’s pathways are not comprehensive NQF L5-10 pathways. (Currently, the B Eng degree at NQF Level 8 is the highest qualification that ECSA acknowledge for professional registration. This does not, however, prevent ECSA from adding further pathways for professional registration in the future.) It may be useful for engineering departments to incorporate one or more of ECSA’s pathways into one or more NQF L5-10 pathways in order to prepare students not only for professional registration, but also to provide them with a route via which they can become involved in research (if research outputs are important for a specific department).

Universities and decision-making: programme and qualification capacity, and the creation of more Master’s and PhD learners (DHET, 2013: xiv, 33, 35). It is for these reasons that a full NQF level 5 to 10 pathway is proposed as the PQM for mining engineering at Unisa ➤ To offer multiple ‘comprehensive’ (NQF L5-10) pathways. This may be done if the appropriate conditions exist. This option is discussed in more detail below.

The following provides some information and guidelines on the usefulness, advantages, disadvantages, and applications of various NQF L5-10 pathways and combinations of pathways.

Offering pathway 1 (starting with the B Eng) Pathway 1 consists of three LPs, namely the B Eng, Master’s, and Doctoral degrees (see Figures 1 and 2). It therefore includes the (ECSA) pathway for developing a person with a NSC to the level where the academic requirements for registration as a candidate professional engineer are met.

Pathway 1 is offered by many engineering faculties at traditional academic universities. This pathway has basically not been affected by the HEQSF. Some of these universities may be research-intensive institutions that want to focus on the postgraduate level. One of the advantages of this pathway is that the B Eng degree is a well-established qualification that has been offered by a number of universities for a considerable period of time. Two (potential) disadvantages of this pathway are as follows: ➤ The admission requirements for the B Eng degree at most universities preclude many students from registering for this LP, e.g. symbols for Mathematics and Physical Science. ➤ Offering a B Eng degree by means of distance learning is problematic due to the fact that it is a 560-credit qualification and a student in full-time employment may take very long to complete this degree. Employed students will also have to spent time away from the workplace to attend laboratory sessions and non-workbased WIL if incorporated into such a programme.

Offering pathway 2 (starting with the Dip [Eng Tech]) Pathway 2 consists of five LPs, namely the 240-credit Dip (Eng Tech), Adv Dip (Eng), B Eng Tech Hons, M Eng, PhD, as well as a WIL component (HEQSF, 2013: 24). See Figures 1 and 2. This pathway includes a number of the vocational programmes listed in Table III. It is therefore an appropriate option for a department at a comprehensive university that wants to offer both vocational and academic programmes and that has a history of offering vocational programmes.

Pathway 2 will allow such a department to keep its current student body (since a switch to pathway 1, for example, will result in higher admission requirements (for the B Eng degree) which only a much smaller percentage of its current student body is likely to meet). (This example points to the ‘path-dependency’ factor, which is not discussed in this paper.) One would expect more diversity in the staff profile of a department offering this pathway: some members will need to have gained industrial experience and done vocational

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programmes themselves, while some staff will, in addition, have to be recognized researchers with the ability to supervise research at Master’s and Doctorate levels. One of its main strengths is that this pathway could prepare students for three of ECSA’s four categories of professional registration, namely Certificated Engineer, Technologist, and Technician (Figure 2). It is therefore the single most flexible pathway in terms of ECSA’s requirements for professional registration, and should be of great use to the general student population and, indeed, most industries.

One of the advantages of pathway 2, which is particularly relevant to distance learning, is its multiple exit levels combined with the fact that it includes some of the smallest programmes (in terms of credit value). This is very important, since a learner who is in full-time employment may take twice as long to complete a programme compared with a student at a class-based institution. More exit levels and shorter programmes in a pathway therefore reduce the risk of (1) non-completion on the part of the student, and (2) reduced government subsidy on the part of the university.

An important advantage (from a university’s perspective) of pathway 2 is that WIL does not form a compulsory requirement for the 240-credit diploma, but follows after the diploma has been completed – this means that the 240 credit diploma is a prerequisite for the WIL component for learners aspiring to achieve access to the Advanced Diploma. In terms of the HEQSF, universities will be held responsible for placing students in industry for work-based, WIL (HEQSF, 2013: 11, 24). In 2011, 860 students were registered for the current National Diploma in Mining Engineering at Unisa. The number of students that complete the National Diploma successfully are fewer and more manageable as far as WIL placement is concerned. There are a number of reasons for this relatively low throughput rate, but two are worth mentioning: 1) Distance learning requires the student to be both highly disciplined and independent 2) The admission requirements for mining at Unisa are much lower compared with those of the University of Pretoria and the University of the Witwatersrand.

In pathway 2, an additional 120 credits to the 240-credit Diploma must be completed in order to obtain admission to the level 7 Advanced Diploma in Engineering. This is described in the HEQSF (2013: 25) as follows: ’Candidates who complete the 240-credit Diploma may enter an Advanced Diploma upon successful completion of a work-integrated learning component or a combination of work-integrated learning and coursework equivalent to 120 credits that is approved and accredited by an education provider and/or a professional body and a QC.’ (See Figure 1.)

Offering pathway 3 (starting with the Dip [Eng]) Pathway 3 consists of five LPs, namely the 360-credit Diploma, Adv Dip (Eng), B Eng Tech Hons, M Eng, and PhD. See Figures 1 and 2. Like pathway 2, it includes a number of the vocational programmes, listed in Table III. Pathway 3 may achieve the same basic objective as pathway 2, but includes WIL in the (360-credit) diploma. The The Journal of The Southern African Institute of Mining and Metallurgy

Universities and decision-making: programme and qualification is recommended only for very large, extremely well-resourced departments with very high numbers of students.

disadvantage of pathway 3 is that the lack of sufficient placement opportunities for the 360-credit diploma may prevent some students from graduating in the shortest time possible and thus influence a university’s throughput rate and government subsidy. Pathway 3 – where the diploma includes a WIL component (as indicated in Figure 2) – is recommended only in cases where the university can guarantee WIL placement for students. (See ‘Brief comment regarding WIL placement’.)

General remark regarding the need for a diversity of educational pathways Engineering departments in the country will have to collectively offer a diversity of LPs to ensure that suitable numbers of technicians, technologists, certificated engineers, and professional engineers are educated and trained to cater for the country’s need for scarce skills i.e. range of engineering practitioners.

Offering pathway 4 (starting with the B Eng Tech) Pathway 4 consists of four LPs, namely the B Eng Tech, B Eng Tech Hons, M Eng, and PhD. See Figures 1 and 2.

Pathway 4 is an option for a UoT or CU that elects to implement the new B Eng Tech as an ‘alternative’ to the current B Tech degree/national diploma combination.

(Traditional academic universities are not excluded from this pathway.) One advantage is that this could prepare students for two categories of ECSA professional registration (Table V). Another is that students obtain a ‘degree’ and not a ‘diploma’ at undergraduate level – this is important for some students. (In pathways 2 and 3 students will obtain a ‘degree’ for the first time only at postgraduate level – the honours degree.)

Offering pathways 1 and 2 The combination of these two pathways is an appropriate option for a fairly well-resourced department that has enough students to ensure the viability of both pathways. It combines the strengths of pathway 2 with the more ambitious and elitist pathway 1. This may be a good option for a large department at a comprehensive university. All four levels of professional registration with ECSA can be covered by means of this combination (Table V).

Offering other combinations of pathways Other combinations of pathways are also possible (e.g. offering pathways 1, 2, 3, and 4). Offering all four pathways

Brief comment regarding WIL placement The HEQSF (2013: 11) provides for at least five types of WIL: simulated learning, work-directed theoretical learning, problem-based learning, project-based learning, and workplace-based learning. One type of WIL which is currently being used is workplace-based learning. As far as this type of learning is concerned, the HEQSF (2013: 11) states that: ‘Where the entire WIL component or any part of it takes the form of workplace-based learning, it is the responsibility of institutions that offer programmes requiring credits for such learning to place students into appropriate workplaces.’ The implication of this requirement is different from current practice (at Unisa) where it is also the student’s responsibility to try and find WIL placement, an internship, employment, or a bursary with a mining company. The WIL office at Unisa enables students to upload CVs to a database that is available to employers. Unisa currently provides students with an experiential learning guide and mentor’s guide for WIL and relies on mines, mentors at mines, and heads of department to ensure that students are exposed to the various areas prescribed. The new HEQSF expect universities to guarantee workplace-based learning. For every learner enrolling on the respective programme, given the fact that universities do not own mines or other workplaces, industry’s participation and support for WIL is crucial. It is, however, also important to note that the HEQSF provides for greater diversity of WIL activities. That said, it is extremely

Table V

Individual NQF L5-10 pathways and their relationship with pathways that lead to professional registration with the ECSA Inclusion of the ECSA pathways for registration as professional ...

Engineer

Certificated Engineer

Technologist

Technician

1

Yes

No

No

No

2

No

Yes

Yes

Yes

3

No

Yes

Yes

Yes

4

No

Yes

Yes

No

1 and 2

Yes

Yes

Yes

Yes

-

-

-

-

Yes

Yes

Yes

Yes

...

(Other combinations, e.g. 1, 2, and 3)

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NQF L5-10 pathway

Universities and decision-making: programme and qualification expensive for universities to provide facilities such as mockup tunnels and stopes and pilot plants, and to develop a virtual reality capability and various forms of simulated learning.

If one assumes that industry may still prefer that the major part of WIL must be workplace-based learning, then this raises the question of how universities can guarantee workplace-based placement for students. This is because providing such placements involves scarce industry resources and not all employers in need of qualified persons may have the capacity to provide students with suitable workplacebased support. One option may be to limit student numbers to what industry is capable of committing to. But should universities follow the example of the various faculties of medicine and veterinary sciences, both of which have stringent pre-screening mechanisms? A university that applies such screening mechanisms needs to know how many students can be absorbed by industry and various other employer organizations. Following the example of the faculties of medicine means that engineering departments in universities will need to have suitable screening and selection mechanisms in place to determine which of the applicants will have the best chances of success in their future careers.

Perhaps another solution is for mining companies to reintroduce the well-known and respected Learner Official programmes.

Pathway 2 has the potential to overcome the above problem (at least, to some extent). The 240-credit Dip (Eng Tech), which will include theoretical and laboratory modules but no compulsory work-placed based modules, can act as a ‘screening mechanism’ before successful (but unemployed) students are employed by industry. The Stage 2 structural and mentored professional development of engineering graduates could be designed in such a way as to incorporate WIL.

proposed ’pathway tool’ that will reduce the complexity (by one level) of this decision-making process. Instead of having to decide on the basis of a menu of thirteen qualifications and associated programmes, I propose that decision-making be done from a menu of four main articulated ‘NQF L5-10’ pathways that also include one or more of the ECSA’s pathways leading to professional registration. The proposed ‘NQF L5-10 pathways’ tool is an attempt to move one step closer to the aim of achieving a structured decision-making approach for designing a PQM at departmental level. A holistic evaluation of the strengths and weaknesses of various ‘NQF L5-10’ pathways and combinations of such pathways are required when deciding on a PQM. The mining section at the Department of Electrical and Mining Engineering at Unisa proposes pathway 2 as its PQM for the discipline of Mining Engineering.

References Department: Higher Education and Training (DHET). 2013. White Paper for Post-school Education and Training: Building an Expanded, Effective and Integrated Post-School System. Pretoria.

ECSA. Candidate Engineer, Candidate Certificated Engineer, Candidate Engineering Technologist, Candidate Engineering Technician. http://www.ecsa.ac.za [Accessed 1 Apr. 2014].

QUALITY COUNCIL FOR TRADES AND OCCUPATIONS. [Sa]. Introduction to the Quality Council for Trades and Occupations (QCTO).

HIGHER EDUCATION QUALIFICATIONS FRAMEWORK (HEQF). Higher Education Act, No. 101 of 1997. Government Gazette 303533, Notice 928. 6 October 2007.

HIGHER EDUCATION QUALIFICATIONS SUB FRAMEWORK (HEQSF). 2013 (as revised). http://sun025.sun.ac.za/portal/page/portal/Administrative_Divisions/INB/ Home/New%20Modules/Revised%20HEQSF%20Jan2013%20FINAL.pdf

Conclusion and the way forward

[Accessed 3 Oct. 2013].

Decision-makers in university departments have to consider many factors, requirements, and constraints when deciding on a PQM. In this paper, I have generalized and expanded ECSA’s pathways for professional registration to suggest a

VAN NIEKERK, D. 2013. Engineering qualifications and the Higher Education Qualifications Sub-Framework (HEQSF). Presentation at the Science Campus of Unisa.

◆

Erratum The affiliation for the author A. Heidary Moghadam published in the SAIMM Journal vol. 113, no. 12, pp. 941–945 entitled: ʻA study on the effect of coke particle size on the thermal profile of the sinters produced in Esfahan Steel Company (ESCO)ʼ, by A. Dabbagh*, A. Heidary Moghadam†, S. Naderi*, and M. Hamdi* was incorrectly listed by the author.

The correct affiliation of the author should be Metallurgy Department, Engineering Faculty, Islamic Azad University, Dezful Branch, Dezful, Iran

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Technical Conference and Industry day

Sustaining Zero Harm Emperors Palace, Hotel Casino Convention Resort, Johannesburg

20–21 August 2014—Conference 22 August 2014—Industry day

‘It always seems impossible until it’s done’ Nelson Rolihlahla Mandela

OBJECTIVES WHO SHOULD ATTEND The conference should be of value to: ˙ Safety practitioners ˙ Mine management ˙ Mine health and safety officials ˙ Engineering managers ˙ Underground production supervisors ˙ Surface production supervisors ˙ Environmental scientists ˙ Minimizing of waste ˙ Operations manager ˙ Processing manager ˙ Contractors (mining) ˙ Including mining consultants, suppliers and manufacturers ˙ Education and training ˙ Energy solving projects ˙ Water solving projects ˙ Unions ˙ Academics and students ˙ DMR ˙ Acid mine drainage.

For further information contact: Head of Conferencing Raymond van der Berg, SAIMM, P O Box 61127, Marshalltown 2107 Tel: +27 11 834-1273/7 Fax: +27 11 833-8156 or +27 11 838-5923 E-mail: raymond@saimm.co.za Website: http://www.saimm.co.za

The conference will focus on improving health, safety and the environmental impact in the mining and metallurgy industry and highlight actions to be taken. It will act as a platform for learning and allow people to share ideas on safety, health and the environment.

This conference aims to bring together management, DMR, Chamber of Mines, Unions and health and safety practitioners at all levels from the industry to share best practice and successful strategies for zero harm and a value-based approach to health and safety. It will address the main challenges in the mining industry such as logistics, energy and safety of employees, contractors and the communities.

SUPPORTED BY:

MMMA

SPONSORSHIP Sponsorship opportunities are available.

Companies wishing to sponsor or exhibit should contact the Conference Coordinator.

SPONSORS:

INTERNATIONAL ACTIVITIES 2014 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 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

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 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 1–2 September 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

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

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

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

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

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

23–24 September 2014 — Grade control and reconciliation Moba Hotel, Kitwe, Zambia 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

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INTERNATIONAL ACTIVITIES 2014 (contineud)

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

The Falls Resort, Victoria Falls, Livingstone, Zambia 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 13–17 July 2015—School on Production of Clean Steel Misty Hills Conference Centre, Muldersdrift 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 12–14 August 2015—The Seventh International Heavy Minerals Conference ‘Expanding the horizon’ 12–13 August 2015—Conference 14 August 2015—Technical Visit Sun City, 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 30 September-2 October 2015—WorldGold Conference 2015 30 September–1 October–Conference 2 October 2015–Technical Visits Johannesburg, 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

19–20 November 2014 — Accessing Africa’s Mineral Wealth: Mining Transport Logistics 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

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

20–24 October 2014 — 6th International Platinum Conference Sun City, 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 12 November 2014 — 12th Annual Southern African Student Colloquium Mintek, Randburg 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

2015 14–17 June 2015 — European Metallurgical Conference Dusseldorf, Germany Website: http://www.emc.gdmb.de 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 6–8 July 2015—Copper Cobalt Africa Incorporating The 8th Southern African Base Metals Conference

The Journal of The Southern African Institute of Mining and Metallurgy

12–14 October 2015—Slope Stability 2015: International Symposium on slope stability in open pit mining and civil engineering Cape Town Convention Centre, Cape Town 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 8–13 November 2015—MPES 2015: Twenty Third International Symposium on Mine Planning & Equipment Selection 8–11 November 2015—Conference 12–13 November 2015—Tours and Technical Visits Sandton Convention Centre, Johannesburg, South Africa Contact: Raj Singhal, E-mail: singhal@shaw.ca or E-mail: raymond@saimm.co.za, http://www.saimm.co.za

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29–30 September 2014—SHAPE: 1st International Conference on Solids Handling and Process Engineering University of Pretoria, 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

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 Hall Core Drilling (Pty) Ltd

Royal Bafokeng Platinum

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

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

â&#x2013;˛

xii

MAY 2014

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

The Journal of The Southern African Institute of Mining and Metallurgy

Forthcoming SAIMM events...

IP PONSORSH EXHIBITS/S ng to sponsor ishi Companies w ese t at any of th bi hi ex or d/ an e th t ac cont events should rdinator -o co ce conferen ssible as soon as po

2014

F

◆ SYMPOSIUM 6th South African Rock Engineering Symposium SARES 2014—Creating value through innovative rock engineering 12–14 May 2014, Misty Hills Country Hotel and Conference Centre, Cradle of Humankind ◆ CONFERENCE 27–28 May 2014, 29 May 2014, Misty Hills Country Hotel and Conference Centre, Cradle of Humankind ◆ SEMINAR 26–30 June 2014, ◆ SCHOOL Mine Planning School 15–16 July 2014, Mine Design Lab, Chamber of Mines Building, The University of the Witwatersrand ◆ CONFERENCE 4–5 August 2014, Emperors Palace, Hotel Casino Convention Resort, Johannesbur ◆ CONFERENCE 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 ◆ 1–2 September 2014,

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

◆ SCHOOL 3rd Mineral Project Valuation School 9–11 September 2014, Mine Design Lab, Chamber of Mines Building, The University of the Witwatersrand ◆ CONFERENCE 16–17 September 2014,

Website: http://www.saimm.co.za

速

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