VOLUME 119
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NO. 6
JUNE 2019
2019/06/18 8:09 AM
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The Southern African Institute of Mining and Metallurgy OFFICE BEARERS AND COUNCIL FOR THE 2018/2019 SESSION
PAST PRESIDENTS
Honorary President
* 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)
Mxolisi Mgojo President, Minerals Council South Africa Honorary Vice Presidents Gwede Mantashe Minister of Mineral Resources, South Africa Rob Davies Minister of Trade and Industry, South Africa Mmamoloko Kubayi-Ngubane Minister of Science and Technology, South Africa President A.S. Macfarlane President Elect M.I. Mthenjane Senior Vice President Z. Botha Junior Vice President V.G. Duke Incoming Junior Vice President I.J. Geldenhuys Immediate Past President S. Ndlovu Co-opted to Office Bearers R.T. Jones Honorary Treasurer V.G. Duke Ordinary Members on Council I.J. Geldenhuys C.C. Holtzhausen W.C. Joughin G.R. Lane E. Matinde H. Musiyarira G. Njowa
S.M Rupprecht N. Singh A.G. Smith M.H. Solomon D. Tudor A.T. van Zyl E.J. Walls
Past Presidents Serving on Council N.A. Barcza R.D. Beck J.R. Dixon M. Dworzanowski H.E. James R.T. Jones G.V.R. Landman C. Musingwini
J.L. Porter S.J. Ramokgopa M.H. Rogers D.A.J. Ross-Watt G.L. Smith W.H. van Niekerk R.P.H. Willis
G.R. Lane–TPC Mining Chairperson Z. Botha–TPC Metallurgy Chairperson K.M. Letsoalo–YPC Chairperson G. Dabula–YPC Vice Chairperson Branch Chairpersons Botswana Vacant DRC
S. Maleba
Johannesburg
D.F. Jensen
Namibia
N.M. Namate
Northern Cape
F.C. Nieuwenhuys
Pretoria
R.J. Mostert
Western Cape
L.S. Bbosa
Zambia
D. Muma
Zimbabwe
C. Sadomba
Zululand
C.W. Mienie
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*Deceased * 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) G.L. Smith (2012–2013) M. Dworzanowski (2013–2014) J.L. Porter (2014–2015) R.T. Jones (2015–2016) C. Musingwini (2016–2017) S. Ndlovu (2017–2018)
Honorary Legal Advisers Scop Incorporated Auditors Genesis Chartered Accountants Secretaries The Southern African Institute of Mining and Metallurgy Fifth Floor, Minerals Council South Africa 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
2019/06/24 8:24 AM
Editorial Board R.D. Beck P. den Hoed B. Genc R.T. Jones W.C. Joughin H. Lodewijks C. Musingwini S. Ndlovu J.H. Potgieter N. Rampersad T.R. Stacey M. Tlala
VOLUME 119 NO. 6 JUNE 2019
Contents
Editorial Consultant
Journal Comment: Furnace Tapping 2018 by J.D. Steenkamp . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
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President’s Corner: 25 years on, and it’s election time by A.S. Macfarlane . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
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D. Tudor
The Southern African Institute of Mining and Metallurgy P.O. Box 61127 Marshalltown 2107 Telephone (011) 834-1273/7 Fax (011) 838-5923 E-mail: journal@saimm.co.za
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Copyright© 2018 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.
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FURNACE TAPPING Development, installation, and operation of a full-copper, deep-cooled slag tapblock for a six-in-line copper matte settling electric furnace B.N. Belford, P. Conradie, and T. Mwanza . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
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Metix has introduced a three-piece ‘deep-cooled’ copper slag tapblock into a six in-line copper matte settling electric furnace. This paper covers the full development life-cycle of the tapblock and illustrates the viability of the design based on operational data that was generated and collected over the period of operation. Phase effects in tap-hole flow – a computational modelling study Q.G. Reynolds, J.E. Olsen, M.W. Erwee, and O.F. Oxtoby . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
527
Challenges encountered with the coupling of phenomena such as multiphase flow and porosity into computational fluid dynamics methods for tap-hole modelling are discussed in the context of electric furnaces used for ferromanganese production. Qualitative and quantitative results of computational models of flow through tap-holes are presented and compared as a function of material parameters to determine sensitivity effects, and in response to operational parameters, to determine typical process behaviour over the duration of a tap. Wear analysis of tap-holes at two ferrochromium production furnaces J.D. Steenkamp . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
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During the annual shutdown of two 63 MVA submerged arc furnaces producing ferrochromium in South Africa, hot excavations of the tap-holes allowed for a study of the wear profiles. Subsequent thermodynamic calculations allowed for the quantification of the potential for chemical reaction between refractory and slag or alloy contributing to wear in the tap-hole area. Tap-hole maintenance and management on furnace 10 at Eramet Norway Porsgrunn J.E. Davidsen and M. Honstad . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
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A worn or short tap-hole channel can lead to more violent flow, significantly increasing the fume emissions during tapping and leading to more metal being mixed with and lost to the slag. Over time, a worn tapping channel might cause side channels and damage to the lining around the tap-hole area. This article summarizes some of the experiences obtained from tap-hole maintenance that has been done regularly with electrode paste on a 33 MW SiMn furnace, as well as a change to a coal tar pitch-free binder.
International Advisory Board R. Dimitrakopoulos, McGill University, Canada D. Dreisinger, University of British Columbia, Canada M. Dworzanowski, Consulting Metallurgical Engineer, France E. Esterhuizen, NIOSH Research Organization, USA H. Mitri, McGill University, Canada M.J. Nicol, Murdoch University, Australia E. Topal, Curtin University, Australia D. Vogt, University of Exeter, United Kingdom
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Multiphase flow modelling of lancing of furnace tap-holes: validation of multiphase flow simulated in OpenFOAM® M.W. Erwee, Q.G. Reynolds, J.H. Zietsman, and P.J.A. Bezuidenhout. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1
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A cold-model validation was carried out as part of a study that focused on the effect of lancing on flow inside and around the furnace tap-hole. This required the validation of the multiphase InterFoam multiphase flow solver in OpenFOAM®. The sensitivity of the model to different material properties (viscosity and density in particular) is presented with both cold models and the results. Slag carry-over and the production of clean steel P.C. Pistorius . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1
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The adverse effects of slag carry-over include increased consumption of deoxidizers, phosphorus pick-up by the steel, and increased ladle refractory wear. Infrared imaging at longer wavelengths is an effective non-contact method to detect slag in the tapping stream. Various devices that are used to shut off the tapping stream rapidly are described. Managing the tap-hole life-cycle at five submerged arc furnaces producing silicomanganese at Transalloys J.J. Sutherland and J.P. Gous. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1
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The Transalloys plant operates five open submerged arc furnaces with an annual capacity of 180 000 t of SiMn. Tap-hole maintenance is critical to effective furnace operation, allowing proper drainage of the furnace with minimal operator interference. A review of the submerged arc furnace operation, furnace and tap-hole design, daily tapping operation, and maintenance practices for repairing tap-holes is presented.
PAPERS OF GENERAL INTEREST Effect of reclaimed bauxite on andalusite-based refractory castables for tundish applications E. Matinde, and S.L. Msibi. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1
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The effect of reclaimed bauxite as a replacement for virgin andalusite in andalusite-containing refractory castables for tundish applications was investigated. A number of tests were conducted to evaluate the physical, physicochemical, and thermochemical properties of the formulated products. The study indicated that an attractive potential saving could be made when the reclaimed alumina replacement ratio was increased to 22% by mass. Kell hydrometallurgical extraction of precious and base metals from flotation concentrates – Piloting, engineering, and implementation advances K.S. Liddell, M.D. Adams, L.A. Smith, and B. Muller. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1
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The Kell process is a hydrometallurgical treatment option for the recovery of precious and base metals from flotation concentrates to refined products on site. The process is cyanide-free, while eliminating emissions of sulphur dioxide, arsenic trioxide, and other toxic species. Kell has been tested on PGM, refractory gold-silver, copper-gold, and polymetallic concentrates, showing extraction efficiencies of >95% for value metals, and demonstrated in a nine-week integrated pilot campaign at 1:1000 scale to support a bankable feasibility study for a 110 000 t/a commercial plant treating a UG2Merensky concentrate. Release analysis of coal fines: Evolution of the methodology and critical issues involved L. Sahu, S. Bhattacharya, and S. Dey. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1
595
A procedure called release analysis, considered to be the counterpart of the washability process in density separation, has been developed to evaluate the flotation characteristics of coal fines. This paper presents a review of and a discussion on the chronological evolution of release analysis methodology over the decades, and highlights the critical parameters involved in the procedures.
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he 2nd Furnace Tapping Conference (Furnace Tapping 2018) was held in Skukuza, the main camp of the world-renowned Kruger National Park, from 14 to 17 October 2018. As with the 1st Furnace Tapping Conference (Furnace Tapping 2014), Furnace Tapping 2018 was well supported and attended by 144 delegates from 17 different countries spanning five continents. A problem-based event, the Furnace Tapping conference series was established to serve the pyrometallurgical industry at large, focusing on a very specific challenge: tapping of furnaces. Drawing from various perspectives on the topic, participants represented industry, consultants, service providers, and research institutions. The latter were a welcome addition to the conversation with contributions from Mintek, the Norwegian University of Science and Technology, SINTEF, and Carnegie-Mellon University gratefully acknowledged. These institutions all delivered on the call, issued in the Journal Comment for Furnace Tapping 2014, for ‘strong contributions from research institutions (focusing on the application of CFD modelling with associated assumptions, including properties of materials and validation of these models; studies on taphole clay; and studies on refractory interaction with slag/metal/matte specifically under the tap-hole conditions)’. Professor Chris Pistorius, co-director of the Centre for Iron and Steelmaking Research at the Carnegie Mellon University in Pittsburgh, delivered the opening keynote presentation. He discussed infrared imaging at longer wavelengths as an effective non-contact method to detect slag in the tapping stream during the steelmaking process, as well as various devices used to rapidly shut off the tapping stream. Many of these technologies should be transferable to other commodities. Members of the audience expressed their appreciation for the knowledge he shared from a commodity not otherwise presented at the conference. Kobus Sutherland, operations manager at Transalloys, shared his perspective on managing the tap-hole life-cycle as part of his keynote address on the second day of the event. Using the analogy of an ant falling into the trap of an antlion, he explained that the strategy followed at the largest producer of silicomanganese in Africa is one of staying out of trouble rather than trying to fix things. The practical nature of his presentation was well received by the audience. Over the two days, presentations ranged from operational practices to tap-hole designs, material and equipment selection, and the application of various modelling techniques in optimizing these aspects. The technical programme closed with an interactive workshop on the second day that allowed delegates to reflect on some critical aspects of furnace tapping, steering the direction for potential multidisciplinary research and other collaborations in the field. A well-supported programme of post-conference technical tours was arranged on the third day. Delegates were offered the opportunity to join the industrial tours to the Polokwane Smelter of Anglo Platinum or to visit the facilities of Neven Matthews in Emalahleni. We are grateful for the opportunities provided by both companies. The Kruger National Park, with nearly 2 million hectares of unrivalled diversity of life forms, offers a wildlife experience that ranks with the best in Africa. Apart from animals in the environs of the camp, delegates viewed wildlife while being transported in open vehicles to the welcoming function as well as the conference dinner. Both events occurred in the bush with no fencing between people and animals. The conference dinner was a magical experience, with hyenas patrolling the perimeter and delegates being appreciative of and even participating in the cultural programme of dancing and singing. Once the formal programme for the evening concluded, the Chileans and South Africans even held an impromptu sing-off contest while waiting for the vehicles to depart for camp! The peer review process was taken very seriously. All the papers published in the conference proceedings were independently peer-reviewed. The organizers drew on the expertise and insights of a group of specialists from around the world who generously offered constructive criticism and suggestions. We are grateful for the inputs from these expert reviewers, and also for Nazli Mamdoo, publications coordinator of the SAIMM, who co-ordinated the process and to Antony Cowey who proofread and edited the proceedings. We appreciate the time and effort that authors put into their papers and presentations. The SAIMM has a policy of making papers from the Furnace Tapping conference series freely available via open access: papers for Furnace Tapping 2014 are available at www.saimm.org.za/Conferences/FurnaceTapping/ and for Furnace Tapping 2018 at www.saimm.co.za/Conferences/FurnaceTapping2018/. The assistance of Dr Rodney Jones, with his passion for open access publishing, is gratefully acknowledged. This edition of the Journal contains a selection of the excellent papers that were presented at Furnace Tapping 2018 and proposed for publication in the Journal by both conference peer reviewers. Each of the papers published here has been through a second formal peer review process, this time in line with the requirements of the Journal. I am very grateful to the members of the Organizing Committee and their employers for their contributions to making the event such a success: Hannes Goosen and Vicky Visser (Dango and Dienenthal South Africa), Isabel Nolet (Hatch), Lars Lindstad (Elkem Carbon), Elias Matinde (University of the Witwatersrand), Rodney Hundermark (Anglo American), Harmen Oterdoom (SMS Group), and Quinn Reynolds and Wesley Banda (Mintek). Special thanks are due to the SAIMM conferencing and publishing teams for hosting and ensuring the smooth running of this event. The support of our sponsors and exhibitors is greatly appreciated. I am looking forward to participating in future discussions on furnace tapping, leaving you with words of Oriah Mountain Dreamer in The Invitation: ‘I want to know if you will stand in the centre of the fire with me and not shrink back’.
J.D. Steenkamp Chairperson of the organising committee, Furnace Tapping 2018
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25 years on, and it’s election time
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ecently, I have written several articles around the 25 years of democracy that we have experienced, and the 125 years of the SAIMM’s history. These articles acknowledged the challenges we have experienced as an industry, but focused on the feeling of a ‘new dawn’ for a ‘sunrise industry’. As I write this, we are into election week, and so our minds are focused on the promises and criticisms of politicians jostling for our votes. This makes me think of where we are after 25 years of democracy from a more circumspect standpoint, in order to really unpack what we need to do individually, as an industry, and as an Institute to realize a better future. I think a relevant starting point in this discussion is to understand how life has changed for individuals in our democracy. For those who were relatively wealthy before the advent of democracy, life has probably not changed very much, and in fact is likely to be better, what with the equalization of high-level salaries with their international counterparts, relaxation on foreign exchange restrictions, and globalization. At the other end of the scale, however, are millions of people who still live in poverty, and for whom the promise of a ‘better life for all’ is still a very distant dream. Service delivery protests in places like Alexandra, Soshanguve, and Eersterivier continue, and the frequency of these protests has appeared to increase, from as far back as 1970 according to Steven Friedland from the University of Johannesburg. Meanwhile, in the middle are millions of people who either turn a blind eye to these protests or use their armchairs as a platform to postulate that this is a problem that government should sort out, and that the country is going to the dogs. They may also feel that apathy is fine, as their vote is unlikely to have any influence. At the same time, we have experienced energy shortages, corruption, and state capture, with outcomes from the Zondo Commission continuing to astound us in terms of the sheer scale of interference and capture that took place. The astonishment that we feel is no doubt fuelled by an overwhelming feeling of ‘how could we not have seen this coming?’, thereby creating a kind of collective guilt and a feeling of inadequacy. The much-maligned attitude that energy shortages are the result of the economic success of the country may have a tiny modicum of truth about it, in that the current crisis, fuelled by capture and mismanagement, was also created by poor strategic planning, which should have taken into account the strain that democratization and the resultant influx of people into the land of plenty and urban areas would place on the supply of these and other services. The situation we find ourselves in now as individuals requires us to decide whether we have some collective responsibility to sort this out, or to leave it as a political battleground where the rich get richer and the poor get poorer, and inequality thrives. Ethical leadership that promises a collective and collaborative solution to our future is what I, in particular, am looking for. How does this translate into the mining industry, and what has democracy done for it? The idiom that ‘a system is perfectly designed for the results it gets’ holds true for our situation. Prior to democracy, driven by sanctions and mineral rights ownership and exclusion, our mining industry was populated by large, institutional mining houses. After democracy, this had to change, and the Minerals Policy and the Minerals and Petroleum Development Act resulted in a new system of mineral rights ownership which was aimed at allowing new entrants into the industry, and participation by all. While the intent was noble, there were many practical obstacles that stood in the way of potential entrepreneurs, such as access to mineral rights, to finance, and to markets. I am in the privileged position of having assisted a new startup company to become a successful manganese producer, but this was a long and hard journey. However, it stands as an example of what can be done, and should help other new empowerment companies to develop.
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The rise in commodity prices during the ‘supercycle’ created the opportunity for new entrants into the mining industry, both as startup companies as well as foreign entrants, and the unbundling of the institutional mining houses into commodity-specific entities. The development of new projects and operations, and the expansion of existing operations, created the expectation of new jobs, with the influx of large numbers of people into previously rural and agricultural communities, such as in Sekhukhuneland. The influx of people usually far outweighed the number of job opportunities available, and often, specialist workers such as rock drill operators were still sourced from traditional labour-sending areas as migrant labour. The result has been the creation of highly volatile, complex, and fractious communities in mining areas, often plagued by inappropriate housing policies, crime, social degradation, and poor service delivery. Such a situation has been worsened by the effects of the end of the supercycle and the drop in prices, resulting in mine closures and retrenchments and accompanying anger amongst community dwellers. The memory of the events of Marikana and the loss of life there will remain in our consciousness for ever. Rapidly increasing pressures regarding the ‘social licence to operate’, the protection of the environment, and health and safety matters are now front and centre instead of being peripheral issues to deal with, changing the way companies operate. Once these pressures were realized, legislation was enacted, the Mining Charter defined requirements, and mining companies spent their efforts on ensuring compliance. Thus, in terms of housing, companies looked either to granting their employees housing allowances, or building houses and schools, offering low-interest loans. While these developments have been impressive in many cases, they have created new social problems that were unforeseen, and people have become indebted beyond their means as a result of unexpected financial burdens. The realization now is that the matter requires more than mere compliance, and instead needs strategic dialogue to be able to engage meaningfully, to understand the concerns and expectations of these communities, and society as a whole. This is an imperative, but does place an extra financial burden on companies; so if this is to work, there has to be a set of compacts that create platforms for dialogue and collective action, and a willingness to spread the financial burden. Where this impacts on service delivery, then it needs to address the issues that exist at a local, municipal, provincial, and national level. This requires recognition of the fact that many municipalities, and even provincial administrations, are in financial distress, that they do not have the competence or capability required, and that it is unreasonable or impossible to assume that these administrations can, on their own, solve the issues at hand. Public-private collaboration is essential, but not as charity. Instead, lines need to be drawn in terms of roles and responsibilities in such partnerships, to ensure sustainability through capacity development and the development of partnerships that are beneficial to all stakeholders. I recently visited a number of villages in remote areas of the Northern Cape where, in discussions with community leaders, it became apparent that they felt their specific and gut-wrenching needs, such as accessible health services, water supply, and clean sanitation were not being adequately addressed, either by mining company Social and Labour Plans or by community trust share schemes. Often, in their opinion, these were focused on municipal needs (maybe driven by politicians) or by the needs within complex and partisan community dynamics and community leadership. Coming back to my idiom that a system is perfectly designed for the results that it gets, the fact that we have a fragmented and ineffective landscape in terms of the benefits and impacts that mining has on the social landscape is symptomatic of the way that we have addressed the issue. A fragmented approach has resulted in a fragmented outcome. And what of the role of our stakeholders? Organized labour maintains its position as being there to represent the interests of its members, but in a fragmented landscape of representation where demands have resulted in the longest strikes the industry has seen during these democratic times. These strikes have delivered very marginal benefits to members but created great economic hardship for people and communities. NGOs have become increasingly active, using topical issues and situations for their advantage, to move their agendas forward. None of this fragmented approach has been helpful in the greater scheme of things. So whose responsibility is it to realize the goals of the National Development Plan, and of society as a whole?
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The answer is … everyone’s, as a collective. This can be construed as a naïve assessment, unless it includes a realization that each stakeholder’s mandates should be respected, and that dialogues must take place in a transparent and honest way. Let us understand how some of the popular words and ideas can be turned into reality, so that our industry can indeed realize a new dawn. Modernization of our industry is inevitable, desirable, and necessary. Stakeholders should not shy away from the notion, but instead come together to create a common vision of what modernization means, and what impacts it must have on our employees and our communities. This must be a future that we engineer for the good of all, not one that is created in the interests of one or two alone. This requires dialogue and open engagement. Part of this vision must involve a realization that old ways of working will become redundant, and that the new future must include upgrading of skills to new norms and to meet new needs, whether this be at operator or manager level. It also requires that we understand collectively the role of the industry in the greater society, and the contribution it makes to that society, now and in the future. This includes understanding and stimulating the full value chain and life-cycle of mining, which then leads to the establishment of local industrialization, appropriate beneficiation, and local agricultural projects. These need to be based on dialogue with communities to ensure that the interventions are value-adding for all involved, and not seen as some form of tokenism or handout. Collaboration is a word that is topical but not always popular. We come from a situation where collaboration was seen as stifling competitiveness, and regarded by some as being slow and bureaucratic. The challenge to stakeholders is to find ways by which such bureaucratic slowness and inefficiency can be removed, so that processes are slick and efficient. We find ourselves now in a situation where solutions that are appropriate to all stakeholders must be developed in a collaborative way, and in a way that is fully inclusive of all needs. This then starts to drive the notion of ‘co-creation’ of solutions for the future. Such co-creation must not be undermined in any way by suspicion, mistrust, or lack of transparency and openness. Inclusivity, transformation, and diversity must be celebrated and seen as cornerstones of our collective engagement. All of these aspects are part of a new interpretation of the sustainability of the industry, which is considerably more complex than the three pillars of sustainability as espoused in 2002, being economic success, environmental care, and social responsibility. Indeed, the new view of sustainability inherently accepts the fact that mining is limiting, because it relies on a finite resource. Instead, mining and processing must recognize the need to develop successful post-mining landscapes, where the mining legacy is not one of destruction, exploitation, and profiteering, but rather one which co-creates new industries, new skills, sustainable communities, and a future that we can all strive for. The Institute is in no way immune to these changes. It must transform from a purely technical organization that satisfies the needs of technical professionals and academics to an organization that actively provides the opportunity for dialogue and discussion on all topics and issues that will create the minerals industry of the future, as a pillar and foundation for a successful, transformed, and inclusive South Africa. Your Council and Office Bearers will continue to assess and reassess the strategic direction of the SAIMM to support this.
A.S. Macfarlane President, SAIMM
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Metix (Pty) Ltd, as part of the SMS Group, is a leading global provider of process plants and engineering services for the pyrometallurgical industry spanning both the ferrous and non-ferrous sectors. Metix operates as a standalone business unit in South Africa from its headquarters in Johannesburg. The company combines more than 100 years of experience with the knowhow of more than 700 smelter projects worldwide.
Our EXPERTISE Metix provides process solutions to the following industries: Ferrous
Ferrochrome Ferromanganese Silicon Manganese Ferrosilicon Ferrovanadium Ferronickel Ferroniobium Pig iron
Non-ferrous
Copper matte Platinum matte Calcium Carbide Ilmenite Silicon metal Precious metals High Purity Alumina Fused Magnesia
Met
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E E E C C S F E C
Metix core technologies for the above include: – – – – – – – – – –
AC Furnaces (circular / rectangular, 3 to 135 MVA) DC Furnaces (circular, 3 to 72 MW) Submerged, brush and open arc operation Gas cleaning (wet / dry scrubbing and baghouses) Energy recovery and cogeneration (open / closed) Medium and low carbon converters Slag cleaning HV / MV / LV regulation and distribution SCADA control & HMI Refining converters
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Ancillary systems designed by Metix include:
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Raw material and product handling Agglomeration, drying and sintering Water treatment and de-watering Cooling water and air Dust extraction and fume abatement Slag granulation Pig casting machines Services and utilities (water, air, gases) Plant infrastructure
The Journal of the Southern African Institute of Mining and Metallurgy
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ding ices rous
outh pany now-
ies:
PASSION FOR METALS
Metix project formats include:
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Engineering, Procurement and Construction (EPC / LSTK) Engineering, Procurement and Supply (EPS) Engineering, Procurement and Construction Management (EPCM) Construction supervision Commissioning and plant optimisation Services (warehousing, repairs, maintenance, technical services) Feasibility studies Export financing Consulting assignments
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5yr Gaurantee on new Metix lower electrode
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Polymet Solutions partnership for ILTEC, Urbangold and extended non-ferrous solutions
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Co-generation solutions for open and closed furnaces
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SMS converters for low and medium carbon ferroalloys plants
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Modern new offices
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Metix (Pty) Ltd. South Africa 204 Rivonia Road, Morningside, Sandton Tel: +27 (11) 676 2300 Fax: +27 (11) 676 2323 E-Mail: info@Metix.co.za www.metix.co.za
2019/06/24 8:24 AM
The Southern African Institute of Mining and Metallurgy
In April 2019, WorleyParsons and Jacobs’ Energy, Chemicals and
is proud to present the
Resources division, two global leaders in engineering, technical and professional services, came together under a new name, Worley. Our customers now
16th Annual Student Colloquium
Embracing Technology and Innovation in the Minerals Industry
have access to a wide network of highly capable people around the globe to help them meet
Date: 16 October 2019 Venue: Johannesburg
the world’s changing energy, chemicals and resources needs.
The Southern African Institute of Mining and Metallurgy has been organizing and presenting the annual Student Colloquium since 2002, to afford the best final-year mining and metallurgical students an opportunity to present their final year projects to an audience of mining and metallurgical industry experts. These students are our future young professionals and will be fundamentally affected by how the industry operates. We have to support and assist our future young professionals! As Nelson Mandela observed: ‘Education is the most powerful weapon which you can use to change the world’. The SAIMM cordially invites our experts in the field to meet the fine calibre of young professionals who are about to embark on their careers in industry. There are 11 mining and 11 metallurgical presentations planned for the event, to be held at Johannesburg on 16 October 2019 . The top five in each discipline will have the opportunity to be published in the prestigious SAIMM Journal in April 2020. The presentations selected will be required to be submitted in the form of draft papers before 8 October 2019.
Our strategy is: To contribute to the nurturing of prosperous and empowered young professionals. FOR FURTHER INFORMATION CONTACT: Yolanda Ndimande | Conference Co-ordinator SAIMM | E-mail: yolanda@saimm.co.za Tel: +27 11 834 1273/7 | Website: www.saimm.co.za
worley.com
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Mining Ad 297x105.indd 1
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Development, installation, and operation of a full-copper, deep-cooled slag tapblock for a six-in-line copper matte settling electric furnace B.N. Belford1, P. Conradie1, and T. Mwanza2 Affiliation: 1 Metix (SMS Group), South Africa. 2 First Quantum Minerals’ Kansanshi Mining Plc, Solwezi, Zambia. Correspondence to: B. Belford
Email:
brett@metix.co.za
Dates:
Received: 11 Mar. 2019 Revised: 8 May 2019 Accepted: 10 May 2019 Published: June 2019
How to cite:
Belford, B.N., Conradie, P., and Mwanza, T. Development, installation, and operation of a full-copper, deep-cooled slag tapblock for a six-in-line copper matte settling electric furnace. The Southern African Insitute of Mining and Metallurgy DOI ID: http://dx.doi.org/10.17159/24119717/663/2019
Synopsis Metix has introduced a three-piece ’deep-cooled’ copper slag tapblock into a six in-line copper matte settling electric furnace. The tapblock was designed without any internal refractory; it relies purely on the development and maintenance of a freeze lining on its hot face and within the tapping channel to protect the copper and promote the integrity of the surrounding endwall refractory lining. In addition, the geometry of the tapblock further supports the sidewall refractories, reduces the impact of funnelling and, therefore, the need for premature rebuilding of the furnace endwall, and facilitates easy removal and replacement of the wear components even under ’hot’ furnace conditions. Eighteen thermocouples embedded in the copper components ensure full resolution of the copper condition, enabling plant personnel to identify wear rates, estimate freeze lining thickness and stability, and schedule replacement of spares and wear components well in advance. At the time of writing the tapblock had been in operation for 6 months with approximately 350 000 t of slag having been tapped through it. Every indication is that the design will limit the need for premature relining of the furnace endwall due to rapid deterioration of the tapblock and its surrounding refractories. Keywords hot repair, slag, tap-hole, tapblock, refractory, condition monitoring.
Introduction Nonferrous furnaces rely heavily on the maximization of refractory campaign life to increase profitability and reduce planned downtime. In large electric copper slag-cleaning furnaces utilizing insulating linings, campaign lives of six or more years are realistic if consideration is given to each aspect that can degrade the refractory lining during the campaign. An area that is particularly vulnerable is the slag tap-hole and the local surrounding refractory. To avoid premature relining of the furnace walls, every effort should be made to consider not only the tap-hole design itself, but also the selection of the sidewall refractory material and its construction immediately adjacent to the tap-hole, and furthermore the interplay between the tapblock and this adjacent refractory. A holistic approach to the design that considers these interdependent features in the slag tapping area can reduce the incidence of accelerated refractory wear (‘funnelling’), support the surrounding refractories to prevent their collapse, introduce easily replaceable wear components that can be exchanged with the furnace still ‘hot’, and encourage and maintain a stable freeze lining at the refractory hot face that maximizes the tapblock and refractory lifetime. With the above in mind, this paper will discuss the development of a copper tapblock for operation in a copper matte settling electric furnace (MSEF) over the full development life-cycle, with the aim to prove the viability of the tapblock based on real operational data collected over the period of its operation.
Background Metix (SMS Group) supplied a 24 MVA rectangular MSEF to First Quantum Minerals (FQM) Kansanshi Mining plc (KMP) for construction of their smelter in 2013/2014 (Chikashi et al., 2016) (Figure 1). The The Journal of the Southern African Institute of Mining and Metallurgy
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Development, installation, and operation of a full-copper, deep-cooled slag tapblock furnace internal dimensions are 24.5 m length × 7.6 m width × 5.45 m height. It uses three 8 MVA single-phase transformers to provide power to six 1.2 m Söderberg electrodes employing Metix upper and lower electrode technology. Principal feed to the furnace is through the sidewall from an Isasmelt furnace which receives dry concentrate at a rate of 161 t/h with a turn-up to 200 t/h. Ten charging bins and their associated vibratory feeders and slide gates introduce a mixture of coke and limestone into the furnace. Two additional launders through the furnace endwall allow for the reintroduction of Peirce-Smith converter (PSC) slag to the furnace as well (Figure 2). The furnace employs an insulating sidewall lining of fused magnesia-chromite bricks, 400 mm thick, separated from a water-cooled sidewall panel by a 50 mm thick conductive ramming layer. The working lining and sub-hearth employ magnesia-chromite (unfused) bricks arranged as an inverted arch and butting up against skewback bricks against the sidewalls. The hearth is cooled by means of ducted cooling air supplied by 14 axial fans along the length of a grillage beam assembly. The furnace roof utilizes a suspended refractory brick system supported by a network of carrier beams off the nine furnace frames that house the shell and limit its transverse expansion. Unlike the sidewalls, the endwalls are restrained by a tie-rod and cup spring arrangement that offers controlled compression of the working lining and sidewall refractories while still allowing for the unavoidable expansion over the length of the furnace. The MSEF is provided with the following feed (approximate quantities):
Figure 1—Night photograph of the KMP smelter. The tallest structure at the top left houses the Isasmelt furnace, with the MSEF located to the right on the ground level. The existing KMP mining and processing sites are seen in the background
Figure 2—Rendered 3D image of the complete Metix scope for the KMP MSEF supplied in 2013
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➤ Matte from Isasmelt ➤ Slag from Isasmelt ➤ Return slag from PSC ➤ Metallurgical coke ➤ Limestone
1 830 t/d 3 100 t/d 700 t/d 22 t/d 25 t/d
With the above as inputs, the approximate output from the MSEF is: ➤ Matte (74% CuS, 21% FeS, and 3.5% Fe3O4) ➤ Slag (0.7% Cu)
1 885 t/d 3 745 t/d
Eight matte tap-holes are arranged along the furnace sidewalls; three on the northern wall and five on the southern wall. Matte is tapped into 50 t matte ladles resting on ladle transfer cars and hauled by winch and rail into the converter aisle for handling by 100 t overhead cranes. Two slag tap-holes are installed on the western endwall opposite the Isasmelt and PSC launders which feed the furnace on the east. Slag is tapped via forged copper launders into 50 t slag ladles and transported by Kress hauler to a remote slag dump. The slag tap-holes are opened and closed during tapping campaigns over the course of a day. Tapping campaigns may last several hours, during which time the tap-holes are opened and closed every 15 to 30 minutes as the slag ladles become full and require removal and replacement. On average, each of the two tap-holes is ‘tapped’ between 35 and 45 times on a given day, with approximately 1 870 t/d slag passing through each tap-hole. Over a year, this equates to approximately 680 000 t slag per tap-hole.
Existing tap-hole design and associated challenges Metix supplied a first-generation slag tapblock (Figure 3) to the KMP MSEF as part of the original engineering, procurement, and supply contract in 2013. This slag tapblock was a two-part, shallow-cooled assembly employing a full copper hot-forged front block and rear hot-forged window block with internal fused magnesia-chromite refractory called the back block extension and tapping bricks, highlighted in Figure 4. The design was unique in that it allowed for the rotation of the front block to change the height of the tapping channel by 200 mm should the operation demand it. After 18 months of operation, during a routine inspection it was found that the hot-face end of the tapping channel had enlarged significantly with a typical bell-mouth appearance. An internal investigation of the tapblock (with the furnace still hot) showed that the internal tapping bricks in the back block had deteriorated appreciably. Wear on the upper extension brick in the rear block, as shown in Figures 5 and 6, was evidence that the bell-mouth wear profile was large enough to start eroding the base of this brick some 400 mm from the hot face and about 200 mm above the tapping channel centreline.
Figure 3—General arrangement assembly views of the first-generation slag tapblock supplied to the KMP MSEF in 2013 The Journal of the Southern African Institute of Mining and Metallurgy
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Development, installation, and operation of a full-copper, deep-cooled slag tapblock Improved slag tapblock design The design mandate Upon discovery of the accelerated wear of the refractories surrounding the existing tapblock, a project was initiated to consider design changes that would achieve improvements on several fronts. These included:
Figure 4—Refractory brick arrangement and naming convention of the first-generation slag tapblock supplied to the KMP MSEF in 2013. Back block extension and tapping bricks highlighted in purple
Figure 5— (Left) View of slag tap-hole hot face wear profile. Bell-mouht size estimated to be 650 x 400 mm. (Right) Photo through the rear copper tapblock from outside the furnace showing the loss of bricks in front of the tap-hole and the deterioration of the tapping bricks in the back block
The bell-mouth hot-face opening was estimated to be 650 mm high and 400 mm wide. The hot-face tapping bricks and the endwall bricks surrounding the tapping channel had been largely lost to wear (Figure 6), and a concern developed that the rows of refractory bricks above the tap-hole would start to collapse as the support beneath them was lost. The original design had not employed a lintel block above the tap-hole. Given the shallow nature of the copper tapblock and the loss of the supporting refractories immediately in front of it, the operations team was unable to replace the hot-face tapping bricks and arrest the ‘funnelling’ of the surrounding refractory. Owing to the design, only the back block tapping bricks could be replaced The accelerated wear around the slag tapping channel was not anticipated and was unwelcome given that the targeted campaign life for the first refractory lining was at least four years. After detection of the wear problem, the KMP operations team performed ongoing measurements of the residual refractory lining thickness and a sidewall and endwall relining was planned for 30 months into the furnace’s first campaign. In comparison, the copper components of the tapblock were found to be in very good condition at the 18-month (and subsequent) inspections. The Journal of the Southern African Institute of Mining and Metallurgy
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➤ Designing the tapblock such that refractories above the tapping channel were fully supported by the block without the need for a lintel ➤ Eliminating refractory from the copper blocks ➤ Instituting ‘deeper’ copper cooling to achieve a freeze lining at the hot face to reduce, or preferably neutralize, funnelling ➤ Incorporating re-usable thermowells that could house conventional thermocouples for condition monitoring of the copper ➤ Enabling hot repair of the tapblock by designing the wear components to be replaceable within 12 hours ➤ Ensuring that any ‘permanent’ components of the tapblock employed redundant cooling, i.e. two cooling channels with the possibility to isolate one channel in the event of damage and continue to run on the standby channel.
Design considerations and final design Metix carried out a project on behalf of KMP to investigate the incorporation of the improvements listed in the previous section in a new slag tapblock for the MSEF. At the outset of the investigation, several constraints were identified. The new tapblocks were replacing the existing two, and there was little incentive to modify the steel water-cooled endwall to accommodate the new design. The shell endwall comprises closed-circuit serpentine cooling which, if modified during a shutdown, presents challenges and risks. The request was to try to accommodate the existing opening size and means of attachment to the shell endwall with the new design. The existing block incorporated only two cooling circuits with a flow rate of 15 m3/h to each circuit. The new block was envisaged to have three copper parts, with one of the parts employing a redundant (standby) circuit. This would mean at least two additional circuits per tap-hole (four in total) and an additional water supply of 30 m3/h per tapblock. The final tapblock design incorporated three copper parts; the inner block, the protective sleeve, and the front block. The assembly of the three blocks is shown in Figure 7.
Figure 6—(Left) Endwall refractory wear in the vicinity of the tapblock observed at the time of refractory relining. (Right) Photograph through the rear copper tapblock, view angled upwards, showing the hot face wear on the lower-front extension bricks above the tap-hole. Evidence of this wear suggested that the bell-mouth funnel extended all the way back from the front of the refractory to the extension bricks; some 400 mm VOLUME 119
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Development, installation, and operation of a full-copper, deep-cooled slag tapblock
Figure 7—Isometric (left) and vertical section view (right) of the new tapblock design
The final design fulfilled the design intent as follows: ➤ Self-support of refractories above the tapblock As shown in Figure 7, the final tapblock design featured three major copper assemblies; the inner block (a wear part), the protective sleeve (a semi-permanent part designed to last the life of the refractory campaign), and the front block (the outermost block, comparable to a ‘mickey/monkey’ block in other designs). The protective sleeve formed the outermost copper assembly and was designed with the express intent of surviving the refractory campaign, the target for which was six years. To do so, the hot face was retracted behind the face of the inner block; the intent being to protect it with a freeze lining after the deterioration of the refractory immediately in front of and surrounding the block. The rectilinear shape of the protective sleeve and the secure method of attachment to the steel shell guarantee that the protective sleeve can support the above and adjacent refractories of the sidewall. This feature ensures that regardless of refractory wear, a collapse of sidewall refractories in the vicinity of the tapblock is unlikely. ➤ E limination of refractory from the copper blocks Extensive computational flow and heat transfer calculations were undertaken to design a tapping channel that was matched with adequate cooling to achieve a stable freeze lining in the channel (a thorough description of the tapblock modelling is discussed in the next section). The thermal modelling predicted a tapping channel freeze lining thickness of between 30 and 50 mm, which was well supported by on-site measurements after four months in operation (refer to Figure 15). With the ability to ensure a stable freeze lining, the refractory could be removed from the tapping channel because adequate protection of the copper, as well as thermal insulation, was provided by the frozen slag. ➤ D eeper cooling to neutralize funnelling The original tapblock protruded only 50 mm past the inner face of the shell, i.e. only into the conductive ramming layer behind the 400 mm thick mag-chrome refractory bricks. The new tapblock was considerably deeper, with the protective sleeve terminating 180 mm from the front of the bricks (270 mm from the sidewall inner face) and the inner block terminating 100 mm from the front of the bricks (350 mm from the endwall inner face) The front or ‘hot face’ of the protective sleeve was designed to work in conjunction with the inner block to form and support a freeze lining in close proximity to the molten bath.
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Figure 7 shows that the inner block projects 80 mm beyond the front of the protective sleeve. In addition, the protective sleeve employs dovetails in its front face that serve to bind it to the solidified material layer immediately in front of it. The projection of the inner block was designed to form a shelf to support the freeze lining. The cooling intensity of both the protective sleeve and the inner block is adequate to freeze the copper slag and to increase the thickness thereof to a point where it shrouds both the front of the protective sleeve and the ‘bullnose’ profile of the projecting inner block. Following an incident of suspected lancing damage, inspections carried out six months after the slag tapblocks went in to operation indicated an acceptable level of funnelling for the period (refer Figure 8). Evidence that funnelling is limited or neutralized over the full campaign life will be possible to obtain only after inspecting the tap-holes after three years. Inspection at the same six-month inspection interval showed little to no wear on the front of the inner block (Figure 8). Photographs taken from outside the furnace (Figure 9) show the freeze lining mass clearly when the inner block is removed. Both findings confirm that the freeze lining is stable (i.e. not eroded away during tapping) and somewhat permanent.
Enabling ‘hot repair’ of the tapblock As mentioned before, after six months of operation the inner slag tapblock was removed for inspection following an incident of suspected lancing damage. The inner block was removed with the furnace still hot by undertaking the following steps. - The furnace was tapped to ensure that the liquid slag level was well below the slag tap-hole level - Power to the furnace was turned off
Figure 8 – (Left) View of ’hot’ face of inner block after 6 months in operation. Minimal to no erosive wear was observed. The damage evident in the photograph is thought to be from lancing and not material erosion during tapping. (Right) View of slag tap-hole inside the furnace after 6 months in operation. Limited funnelling observed for the period
Figure 9—Photograph looking down though the protective sleeve void after the inner block has been removed. The jagged grey-brown mass at the far end is the freeze lining that has remained in place after the withdrawal of the inner block The Journal of the Southern African Institute of Mining and Metallurgy
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Development, installation, and operation of a full-copper, deep-cooled slag tapblock - The castable material and splash plates adjacent to the tap-hole were removed - The thermoelements were removed from the front and inner blocks - The front block was removed with conventional rigging - Use jacking locations on the inner block, fasteners were inserted and tightened to eject the inner block - A purpose-built trolley on a crawl beam was attached to the front of the inner block and supported the weight of the inner block as the trolley was pulled backwards to remove it - To reinstall the inner block the process was reversed - The protective sleeve was not changed. The above removal and reinstallation process of the inner block took less than the allocated 12 hours.
Redundant cooling on permanent components A feature of major importance was the adoption of redundancy on the ‘permanent’ components of the tapblock. As indicated in Figure 7, the protective sleeve was deemed to be a semipermanent feature of the tapblock with the intent that it survives at least the targeted six-year campaign life. Given the protective sleeve’s role in supporting adjacent refractories, removal and replacement of this component would require access into the furnace. To reduce the likelihood of needing to replace this block, two cooling coils of Monel 400 construction were adopted within the copper body forming the protective sleeve (refer Figure 10). Each cooling coil was designed with its own water inlet and outlet and a dedicated circuit to supply 15 m3/h per coil during normal operation In the event of damage to the protective sleeve to the extent that the frontmost cooling coil is compromised (highlighted red in Figure 10), this circuit can be isolated and the secondary coil (highlighted yellow in Figure 10) can continue to fulfil the cooling needs of the protective sleeve. Even with only a single coil in operation, the copper protective sleeve is designed to be adequately cooled such that replacement during the refractory campaign period is not required. At the time of writing, no noticeable damage had yet been inflicted on the protective sleeve.
Computational flow and heat transfer modelling The ‘Holy Grail’ of thermo-flow modelling of a slag system would ideally be a fully defined set of all temperature-dependent material properties across the entire operating range, i.e. from
Figure 10— Simplified transparent 3D isometric view of the tapblock protective sleeve showing the two independent cooling coils (secondary coil highlighted in yellow, and front in red The Journal of the Southern African Institute of Mining and Metallurgy
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solidified slag, through the transitionary ‘mushy’ phase, into superheated liquid. This will remain daunting for designers for two main reasons. ➤ Firstly, the experimental determination of the transitionary and superheated liquid material properties at temperatures exceeding 1000°C is difficult. ➤ Secondly, the transitionary phase exhibits nonlinear behaviour, as the slag constituents with higher solidus thresholds freeze preferentially. Analytical tools like FactSage, based on very large databases of empirical data, are invaluable in the pursuit of this ideal, but hinge on the accuracy of the simulated models, the information in the database, and the accuracy of the slag composition information as supplied by the user. Even for copper slag systems, where the macro-behaviour and chemistry are well understood, reported values for specific properties of frozen slags, such as thermal conductivity, can vary by up to 400% (Nelson and Hundermark, 2014), making accurate dynamic modelling of a slag tap-hole and freeze lining system a challenge, especially as part of an accelerated product development cycle. In a tapblock, the thermal conductivity of the frozen slag lining can be up to 150 times less than that of the copper block. Therefore, the freeze lining thickness will be the single most influential factor (‘throttle’) in the energy balance of the tapblock. Despite this uncertainty in absolute properties, the consensus is that a frozen slag lining is achieved reasonably easily, and once established, remains stable and robust (Nelson and Hundermark, 2014) provided that no significant external influence, for example oxygen lancing, results in a disruption of the freeze lining. If experimentation, analytical modelling, and sensitivity analyses are utilized, temperature-dependent material property accuracies of between 10 and 15% can be compiled. However, this exercise would be time-consuming, especially in the absence of tools like FactSage, making it unfeasible in the context of a product development cycle shorter than three months. For the tapblock under discussion in this paper, it was therefore, necessary to devise an alternative thermo-flow modelling methodology and applicable evaluation metrics.
Modelling methodology and set-up Instead of ascertaining a single, specific freeze lining thickness using temperature-dependent material properties, a methodology was adopted whereby a series of known freeze lining geometries was imposed on the model and the resultant thermal gradients were analysed to determine a range of anticipated steady-state thicknesses. The modelling approach was therefore reduced to the following steps. ➤ Select an operating cooling water flow rate ➤ Prepare geometry of x mm/y mm freeze lining thickness (refer to Figure 11) ➤ Complete a hydrostatically driven CFD thermo-flow model for a given cooling water flow rate and slag temperature ➤ Establish whether the liquidus isotherm is positioned in the freeze lining or in the slag: o If the isotherm is in the freeze lining and close to the surface of the freeze lining, it represents the upper operating thickness of the freeze lining, i.e. the thickest it will be for the given operating conditions VOLUME 119
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Figure 11—Typical computational geometry indicating the frozen slag lining geometry imposed on the model
o If the isotherm is in the slag and close to the slagfreeze lining interface, it represents the lower operating thickness of the freeze lining, i.e. the thinnest it will be for the given operating conditions ➤ Repeat steps 2–4 for various freeze lining thicknesses, x and y, until one of the criteria above is met, then: o Report the outlet water temperature for the respective cooling water circuits for the selected water flow rate o Probe the copper temperature at known thermocouple positions (refer to Figure 20) ➤ Restart at step 1 with a new water flow rate.
feature smooth bends and the micro-level flow behaviour is not of specific interest, the necessity to capture recirculation zones and vortices by explicitly solving the viscous sublayer, as would be prevalent with mitre bends or abrupt contractions or expansions, is diminished, which meant the mesh size and solution time could be reduced by opting for the k-ε two-equation model. The k-ε model was combined with the ‘Enhanced Wall Treatment’ function to make the solution y+-independent by mitigating any inaccuracies at lower y+ values that arise when using the Standard Wall Function (ANSYS, 2017). The physical complexity of the problem was reduced by including only dominant boundary conditions, as summarized in Figure 13. Since the problem is driven by the hydrostatic pressure of the liquid slag, the inlet of the slag is simply defined as a pressure boundary, at the specified level, that allows slag to enter the domain at 1250°C. Water through the respective cooling circuits is defined as a mass flow inlet correlated to a volumetric flow rate between 15 m3/h and 5 m3/h in order to gauge the system sensitivity to water flow rate, and to evaluate the lower
The modelling methodology is subject to the following core assumptions. ➤ Properties at the operating temperatures of the respective materials remain near-constant ➤ The tap-hole diameter through the front tapblock is the same as the diameter of the tap-hole drill ➤ The furnace level remains constant during the course of each analysis. The greater the thermal conductivity of the liquid and frozen slag, the deeper the cooling will penetrate and thus the thicker the freeze lining that will form. A degree of conservatism in the prediction of the freeze lining stability can, therefore, be introduced by assuming the lower thermal conductivity from the range reported in the literature. The full list of material properties used in the design is summarized in Table I. Even though the slag flow is laminar, the aforementioned methodology requires sufficient resolution on either side of the slag-freeze lining interface. To this end, 15 one-millimetre ‘boundary layers’ (or ‘inflation layers’) were defined in both the slag and freeze lining, as shown in Figure 12. In addition, the water circuits were meshed with the requisite number of boundary layers for the realizable k-ε turbulence model. As the cooling water coils for the inner block and protective sleeve
Figure 12—Boundary layers in slag and freeze lining to increase resolution
Figure 13—Dominant boundary conditions for computational model
Table I
Summary of material properties Material Copper Frozen slag Molten slag Refractory Monel 400 Water
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Thermal conductivity (W/mK)
Density (kg/m3)
Viscosity (Pa.s)
Specific heat (J/kg.K)
Reference/comment
381 2.1 2 1.75 22 0.609
8933 4500 3433 3200 8800 996
- - 0.7 - - 0.798 × 10-3
346 1138 1010 874 427 4178
UNS C81100 KMP internal data Nelson and Hundermark (2014), KMP internal data RHI
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Results and interpretation Based on a laboratory analysis of the KMP copper slag, accurate temperature bands were ascertained for the solidus and liquidus temperatures. Since discrete values for each were required in the determination of the freeze lining thickness, the liquidus and solidus temperatures were chosen to be 1160°C and 1059°C respectively for the tapblock design. For the purpose of this qualitative investigation, it was assumed that the slag freeze lining would extend to the liquidus isotherm. Using the liquidus isotherm as an indicator of the freeze lining extent, it is possible to visually establish how close to convergence the iterative geometry methodology is. Figure 14 shows the liquidus isotherm inside the freeze lining, but not sufficiently close to the slag-freeze lining interface, suggesting that this freeze lining is still ‘too thick’. Similarly, the liquidus isotherm would appear in the slag if the freeze lining was ‘too thin’. While the objective was to get the liquidus isotherm and interface to match exactly, a variance of approximately 1 mm was deemed reasonable to proclaim a converged cycle. For the given water flow rates of 15 m3/h (design flow rate), 10 m3/h, and 5 m3/h, it was established that the freeze lining would vary between 28 and 30 mm in thickness. Due to the inner block resting directly on the protective sleeve, as opposed to being grouted in place as with the other three sides, the freeze lining was predicted to be slightly thicker along the bottom of the tap-hole channel. While condition monitoring and maintenance observations will be discussed in detail in the following sections, it is rewarding to note that the freeze lining thickness of approximately 30 mm observed during inspection of the inner block, as shown in Figure 15, aligns well with the predicted values. This confirms the insensitivity of the freeze lining to changing operating conditions, and for that reason, validates the ‘operating envelope’ approach as a simplified method to predict slag tap-hole performance.
Figure 14—Visual comparison of liquidus isotherm and slag-freeze lining interface The Journal of the Southern African Institute of Mining and Metallurgy
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A typical cross-section through the modelled tapblock (Figure 16) illustrates the thermal gradients throughout the computational domain. It is promising that for this example, all the copper operates below 150°C, which is the conservative accepted softening temperature for non-alloyed copper. Even though these components should not typically be externally stressed, the protective sleeve plays a vital role in the refractory sidewall integrity and could, therefore, be considered a structural component. It is standard practice within Metix to remain below this temperature threshold for ‘structural’ copper components. The copper temperature plot shown in Figure 17 is a useful tool for visualizing the heat load distribution on the copper hot face. The combination of a contraction in flow at the tap-hole inlet, the round tapping channel geometry in a square opening, and the distance to the nearest water channel all contribute to the hot face of the inner block being thermally taxed compared to the rest of the tap-hole assembly. In this application, a compromise will be required between cooling channel proximity to the hot face and sufficient copper thickness to act as a physical and thermal protective buffer in a deep-cooled component. Over and above proving the thermal viability of the tapblock, a secondary purpose of the modelling was to determine the copper temperature at discrete locations in the copper components. The final tapblock solution installed at KMP featured 18 thermocouples (four in the protective sleeve, twelve
Figure 15—Freeze lining thickness observed during inspection
Figure 16—Typical full-scale cross-section through tapblock arrangement VOLUME 119
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Figure 17—Example of copper hot face temperature for a 45 mm freeze lining thickness in channel (dimension x) and 100 mm on nose (dimension y)
in the inner block, and two in the front block) positioned at key locations to provide a condition monitoring functionality that is described later in this paper. A typical cross-section through one of the measurement planes, shown in Figure 18, indicates the locations of the monitoring points of the inner block by means of four crosshairs. Plotting the temperatures at each of the 12 measurement points in the inner block for each converged freeze lining thickness cycle not only gives a useful visual indication of what the temperature monitoring alarms could be set to during physical operation, but also allows a quick comparison of the impact of varying cooling water flow rates. Figure 19 plots the copper temperatures for a given freeze lining geometry for three discrete cooling water flow rates. From this it is reasonable to conclude that once the freeze lining is established, it requires very little cooling to maintain, and manipulating the cooling water flow would be an ineffective means to manipulate the freeze lining thickness, and subsequently the slag tapping rates. The ability to create and maintain a freeze lining in a tapblock offers the designer a margin of forgiveness, especially in applications where the slag superheat is limited; less than 90°C in the KMP MSEF operation. As predicted, the freeze lining does indeed act as the dominant factor in the energy balance of the tapblock and renders the system balance somewhat insensitive to changes in the slag and cooling water flow rates. Nonetheless, this development modelling was grounded on a series of simplifications and assumptions that, although valid, render the results an equal blend of qualitative and quantitative findings, with the single most important outcome being the added understanding of the performance response of this novel tapblock design in this application.
Condition monitoring Methods employed The new tapblock supplied by Metix to KMP is essentially a ‘deep-cooled tapblock’, meaning that cooling water is introduced into the endwall refractory region together with the equipment. This design, while proven in industry, must be respected for the associated dangers introduced by cooling water in close proximity to a molten bath. To perform condition monitoring, three monitoring techniques were used, namely cooling water flow rate (in m3/h), cooling water temperatures (both inlet temperature and ΔT in degrees Celsius), and copper temperature (in degrees Celsius). The first two techniques are common on most water-cooled furnace
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Figure 18—Typical cross-section plot through the copper inner block on one of three measurement planes with thermocouple locations shown by the four black crosshairs
Figure 19—Typical cross-section plot through the copper inner block on one of three measurement planes
elements and will not be described further here. To measure the copper temperature, conventional thermocouples were used. The various measured variables were displayed in real time on the supervisory control and data acquisition (SCADA) system and saved to the operational database for daily trending and interpretation.
Layout of the condition monitoring elements These thermocouples were arranged as follows: ➤ Four measuring locations on the hot face of the protective sleeve, as shown by section A-A in Figure 20 ➤ Twelve measuring locations across three planes in the inner block, as shown by sections B-B, C-C, and D-D in Figure 20. Four measurement points were present on each of the three planes ➤ Two measurement locations beneath the tapping channel in the front block (not shown). To allow the thermoelements to reach the desired measurement locations, a blind conduit of stainless steel was precast into the copper body. Thermoelements of known lengths and 1.5 mm in diameter were fed into the conduits during installation so as to reach the prescribed measurement points. The CFD analysis that was undertaken was used as a tool to estimate the copper temperatures at the different measurement locations and to provide KMP with the expected operating ranges, as well as values for alarm set-points for the copper temperature.
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Figure 20—Thermocouple layout in the protective sleeve and inner block of the slag tap-hole
sensors mounted above the slag ladle bay were used to measure the slag levels in the ladles during operation. By calculating the rate of change of the slag levels measured by the sensors, the beginning of a tap and the end of a tap could be clearly identified, and then the associated variables recorded and interpreted. The other product of these calculations was real-time measurements of the slag flow rates (m3/h) during operation. These results were used to validate earlier design modelling work. While all the measured variables were important, particular focus was placed on measuring the hot face temperatures on the protective sleeve (section A-A in Figure 20) and the inner block (section D-D in Figure 20). Each of these locations was monitored by four thermocouples. Each individual thermocouple was recorded and graphed (refer Figure 22), as well as the average of the four thermocouples at each location (refer Figure 23). The average daily trends were then plotted over the course of a month to trend any increases in the copper temperatures (refer Figure 24). Given that up to 50 tapping cycles were recorded per day, each data-point on the individual thermocouple plots (Figure 22) represented the average of the individual thermocouple measurements from the 15- to 30-minute tapping cycle. Standard deviations for each thermocouple during each tapping cycle were plotted to understand the spread of the data during the individual tapping cycle to ensure that temperature spikes were captured.
Usefulness of condition monitoring as the cooling flow meters and thermocouples, was displayed on the SCADA alongside graphics that represented the protective sleeve and inner block components (refer Figure 21). The alarm limits were programmed into the SCADA to alert the operator to unusual operation. The condition monitoring variables were exported daily to an Excel-based analysis tool developed between KMP and Metix. As described in the first section of this paper, each tap-hole may be used up to 45 times in a single day. Each usage cycle would be characterized by drilling (or lancing) the tap-hole open, establishing slag flow until a slag ladle had been filled, and then plugging the tap-hole with the mud gun. Each of these tapping cycles may last between 15 and 30 minutes, during which time a noticeable rise in the copper temperatures and cooling water temperatures would be experienced. Between the tapping cycles, the copper and cooling water temperatures would not be of much relevance as the tap-hole would be closed and deemed to be safeguarded against damage. The analysis tool was thus complicated by having to discard the data between tapping cycles and process only the data-points between the beginning and end of tapping. To this end, level
Over the course of the first seven months of the tapblock operation the trends in Figures 22 and 23 were produced daily and assessed by both KMP and Metix engineers. These trends enabled hotter tapping cycles and lancing events that led to extreme copper temperatures being identified. One useful interpretation of the graphs was to assess whether the tapblock returned to a normal temperature range after a high-temperature event. It was sometimes seen that the copper ran 20°C hotter than the average during certain tapping cycles. After these higher than usual cycles, the tapblock temperatures were monitored to see if the temperatures reduced, thereby indicating that a freeze lining had re-established, and the tapblock was again operating normally. The monthly trends, as shown in Figure 24, were used as a tool to try to identify wear of the inner block and the protective sleeve. While an upward trend on the inner block was discernible in the first four weeks of operation, trends beyond this initial ‘bedding-in’ period were mostly flat. This gave every indication that wear on the inner block and the protective sleeve was not occurring due to an erosive or melting mechanism during tapping.
Figure 21—SCADA representations of the protective sleeve and inner block of the slag tapblock showing the approximate location of the thermoelements and their corresponding instrument tag numbers The Journal of the Southern African Institute of Mining and Metallurgy
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Figure 22—Trend of daily inner block hot face thermocouple measurements. Each data-point is the average of the measurements from the 15- to 30-minute tapping cycle. Standard deviations are plotted to understand the spread of the data during the individual tapping cycles
Figure 23—Average temperature calculated from the four thermocouples at each of the measuring sections in the inner block sections B-B, C-C, and D-D. As can be expected, section D-D, the ’hot face’, is running higher than the others
Conclusions and the way forward This paper has attempted to describe the background to, and the need for, the design and adoption of a modified slag tapblock for First Quantum Minerals’ KMP MSEF. The design mandate for the modified block has been described, together with the operational experience and condition monitoring. During the first refractory campaign of the MSEF, which lasted only 30 months, the refractories above and immediately adjacent to the slag tapblocks were eroded to the point that
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Figure 24—Average temperature calculated from the four thermocouples at each of the measuring sections in the inner block. sections B-B, C-C, and D-D plotted for the month of March 2018. First-order linear trend lines are added to the plot to detect discernible trends in the copper temperatures
refractory collapse was likely if the wall was not relined. With a targeted refractory campaign life of six years, a tapblock capable of limiting the erosive wear and supporting the surrounding refractory was required. KMP employed the services of Metix, the original furnace designer, to investigate the redesign of the slag tapblock to achieve the required campaign life and incorporate other key performance features. The tapblock was designed to fit within the endwall opening of the existing block. The slag tapblock employed three primary The Journal of the Southern African Institute of Mining and Metallurgy
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Development, installation, and operation of a full-copper, deep-cooled slag tapblock components, namely an outer protective sleeve, an unlined inner block, and a front block to act as a flow restriction device. The protective sleeve was deemed to be a ‘permanent’ component capable of surviving the full refractory campaign period. The inner and front blocks were considered wear items and were designed to be replaced within 12 hours with the furnace still ‘hot’. The protective sleeve acted as an integral lintel and flanker cooler by providing support to the refractories above and to the side of the tapblock. To limit the need for protective sleeve replacement, dual cooling coils were adopted to provide redundancy, i.e. in the event that damage was sustained on the protective sleeve hot face, the frontmost coil could be isolated and the secondary coil would be capable of adequately cooling the block. Given the erosive damage witnessed on the refractories of the original tapblock, removal of tapping channel refractories was considered a key feature of the new tapblock. Metix employed extensive CFD modelling of the tapblock to characterize the slag flow rate and the tapblock’s ability to work without the need for internal refractories. The CFD modelling enabled the designers to determine the required cooling to achieve a predicted slag freeze lining thickness. Based on this modelling, the freeze lining stability could be quantitatively estimated and the modelling results predicted that a stable freeze lining would be achieved together with the geometrical features to support it on the ‘hot face’. After 6 months in operation the tapblock was inspected for performance and wear. The inner block was removed to assess the condition of the ‘hot face’ and internal visual inspections of the endwall refractory lining were undertaken to determine the degree of ‘funnelling’. These early indications were positive, with negligible wear found on the front face of the inner block and acceptable refractory wear observed. Further investigations after longer operating periods will be required to evaluate the longerterm performance of the tapblock. Condition monitoring of the tapblock was considered a key feature in the conceptual design phase to provide a real-time tool to evaluate the tapblock condition. Given the adoption of deep-cooled copper components, forewarning of damage that may pose a risk to furnace operation and personnel was required. A total of 18 thermocouples were adapted by precasting blind stainless steel conduits into the copper tapblock bodies. The CFD modelling allowed for the theoretical determination of the copper temperatures at the 18 locations. These theoretical temperatures were used to quantify normal and alarm conditions. The data acquired by the thermocouples, together with other basic variables like cooling water flow rates and water temperatures, was displayed in real time on the SCADA and plotted daily in a Microsoft Excel-based analysis tool. The tool allowed abnormal events to be identified, copper component temperature to be monitored on a tap-by-tap basis, and wear to be trended monthto-month. The inspection carried out after 6 months of operation revealed that the inner block had sustained negligible erosive wear during tapping. There was no evidence to suggest that the protective sleeve had suffered any damage during the period either. A visual inspection of the endwall refractory wear above and adjacent to the tapblock to assess the degree of ’funnelling’ revealed satisfactory results. By all accounts the performance of the modified tapblock had been more than satisfactory over the period. Furthermore, the original inner block life was estimated The Journal of the Southern African Institute of Mining and Metallurgy
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to be between 3 and 6 months before replacement. The intensive cooling and geometrical design features of the tapblock appeared to have generated a freeze lining that was more robust than envisaged during the design. Both Metix and KMP are of the opinion that the inner block lifetimes of 12 months or more are realistic, based on the observations at the 6-month inspection.
Way forward Although no definitive conclusion can be drawn as yet, observations after 12 months of operation indicate that the funnelling of the refractory at the tap-hole inlet has indeed been reduced. KMP will continue to monitor this situation. The expectation is that the funnelling will eventually be arrested by the deep-cooled inner block as the funnel itself will become a stable, self-healing slag freeze lining. Metix and KMP will continue to monitor the inner block closely, with the intention to establish the benchmark for inner block life under normal operational wear and tear. However, the removal of the inner block after just six months due to lancing damage emphasized the risks associated with lancing of a deep-cooled tap-hole. KMP has, therefore, actively pursued the implementation of more controlled and disciplined lancing procedures and techniques, complemented by the use of more advanced lancing equipment and consumables. On the technology side, Metix has already implemented minor optimizations of the inner block geometry to improve the resilience of the component without compromising cooling efficiency, and to also bring the maximum tapping rate in line with KMP’s downstream processes. These design optimizations will be realized with the next spares supply, and in combination with the operational improvements, are intended to reliably extend the life of the inner block beyond 12 months. In order to improve the accuracy of the computational modelling, research and testing will continue, with specific emphasis on deriving a set of temperature-dependent material properties that are accurate enough to provide a reliable engineering solution for the temperature-dependent behaviour of the freeze lining. This endeavour is expected to include further analyses of slag tap-hole samples, continued research, and collaboration with industry partners. This collaboration aims to develop analytical material property models and investigate the incorporation of more dynamic computational models, for example melting and solidification and volume of fluid models, to capture more of the physics present in the tapblock during operation.
Acknowledgements
This paper, and the associated photographs and data, is published with permission of the smelter management team at First Quantum Minerals’ Kansanshi Mining plc. The contributions of our fellow authors and the pragmatic and practical assistance by the operations team are gratefully acknowledged.
References
ANSYS. 2018. Fluent Theory Guide, Chapter 4.16. Near-wall treatments for wall-bounded turbulent flows. Canonsburg, PA. https://www.scribd.com/ document/342817281/ANSYS-Fluent-Theory-Guide Chikashi, H.M., de Vries, D., Hunt, S., Dyussekenov, N., Hanschar, L., and Milovanov, M. 2016. Design, commissioning and operation of a matte settling electric furnace at the Kansanshi copper smelter. Proceedings of the 9th International Copper Conference, Kobe, Japan, 13-16 November 2016. The Mining and Materials Processing Institute of Japan. pp. 642-653. Nelson, L.R. and Hundermark, R.J. 2016. The tap-hole – key to furnace performance. Journal of the Southern African Institute of Mining and Metallurgy, vol. 116. pp. 465-490. u VOLUME 119
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31 July – 1 August 2019 The Canvas Riversands, Fourways Johannesburg, South Africa
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Phase effects in tap-hole flow – a computational modelling study
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Q.G. Reynolds1, J.E. Olsen2, M.W. Erwee1, and O.F. Oxtoby3 Affiliation: 1 Pyrometallurgy Division, Mintek, South Afric. 2 SINTEF Industry, Trondheim, Norway. 3 Aeronautic Systems, CSIR, South Africa. Correspondence to: Q.G. Reynolds
Email:
quinnr@mintek.co.za
Dates:
Received: 15 Mar. 2019 Revised: 13 May 2019 Accepted: 25 May 2019 Published: June 2019
Synopsis The extraction of molten products and wastes through smelting furnace tap-holes during the tapping process is a complex procedure involving many coupled thermophysical effects. Insight into the fluid flow behaviour in such systems can be gained by using computational mechanics tools to build highfidelity models of the fluid flow and other relevant physics. Previous work using such models indicates that the nature of the flow in the tap-hole, as well as operational parameters such as the tapping flow rate, depends strongly on the properties of the material being tapped and the geometry of the tap-hole channel. Additional complications arise when multiple fluid phases (for example, slag and metal) are present simultaneously, when porous coke beds hinder flow to the tap-hole entrance, and when the geometry of the tap-hole changes over time due to cycles of wear and repair. In the present work, challenges with the coupling of phenomena such as multiphase flow and porosity into computational fluid dynamics methods for tap-hole modelling are discussed in the context of electric furnaces used for ferromanganese production. Qualitative and quantitative results of computational models of flow through tap-holes are presented and compared as a function of material parameters to determine sensitivity effects, and in response to operational parameters to determine typical process behaviour over the duration of a tap. Keywords fluid flow, multiphase, modelling, tap-hole, ferromanganese.
How to cite:
Reynolds, Q.G., Olsen, J.E., Erwee, M.W., and. Oxtoby, O.F Phase effects in tap-hole flow – a computational modelling study. The Southern African Insitute of Mining and Metallurgy DOI ID: http://dx.doi.org/10.17159/24119717/671/2019 ORCiD ID: Q.G. Reynolds https://orchid.org/0000-00025196-8586
Introduction Pyrometallurgical production using electric furnaces accounts for substantial portion of the world supply of ferroalloys and other strategic materials (Degel et al., 2015). A range of different furnace designs may be used depending on the metallurgical process, with circular three-electrode submerged arc furnaces (SAFs) powered by alternating current being the most common for commodities such as ferromanganese and silicomanganese (Olsen, Tangstad, and Lindstad, 2007). As shown in Figure 1, manganese furnaces typically operate with a deep charge consisting of an upper layer of loose burden (onto the upper surface of which is fed solid raw material), a molten slag layer partially or completely filled with unreacted carbonaceous reductant (generally metallurgical coke), and a molten metal layer which settles to the bottom of the furnace due to its high density. The solid particles present in the slag and burden layers form porous beds which hinder the fluid flow of molten process material and gases inside the furnace (Olsen, Tangstad, and Lindstad, 2007). In order to remove material from the furnace during normal operations, a single tap-hole is built into the sidewall of the vessel near to the level of the hearth. Tapping involves removing both slag and metal phases through the same tap-hole channel, which makes tap-hole design, operation, and maintenance particularly complex. In the case of processes such as ferromanganese, which operate with no external gas injection or open arcs present at the furnace electrodes, flow of material out of the tap-hole is driven primarily by the hydrostatic pressure generated by the weight of the solid and liquid phases above the tap-hole. Measurement of experimental data from the interior of furnace tap-holes during operation is extremely challenging for a number of reasons, including the high temperatures of the materials involved, physical access limitations, and hazardous environment (see Figure 2). Computational models of tapping can provide useful additional insight in such cases. Full models of manganese smelting furnaces which include the tap-hole region have been produced by several authors, notably Kadkhodabeigi, Tveit, and Johansen (2011) who studied in some detail the effect of furnace conditions on metal, slag, and gas flow from the tap-holes of silicomanganese furnaces
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Phase effects in tap-hole flow – a computational modelling study in porous media, giving the result shown in Equations [1]–[3] for two-phase flow.
[1] [2] [3]
Figure 1—Schematic diagram of the tap-hole region in a ferromanganese furnace showing metal, coke/slag, and burden layers
In these equations, f is the porosity field, ρ is the density field, u is the velocity vector field, P is the pressure field, μ is the viscosity field, g is the interfacial surface tension between the two phases, a is the phase fraction field indicating the volume fraction of slag phase, FP is the drag force acting on the fluid due to the porosity field, and g is the vector of acceleration due to gravity. All field parameters are able to vary in both space and time. The equations must be closed with an appropriate expression for the porous flow drag force. The Ergun equation, a well-known and extensively validated correlation developed for flow in fixed and fluidized beds, was used in this case (Equation [4]). [4]
Figure 2—(l) Lancing a submerged arc furnace tap-hole (image‘“Flying Sparks’ © Joalet Steenkamp 2016), (r) tapping stream detail on a silicomanganese furnace (image ’Follow Me’ © Joalet Steenkamp 2013)
and demonstrated important connections between fluctuations in the furnace interior pressure and tapping flow rates. More detailed models of the tap-hole entry and channel regions are also used in furnace design (Muller, Zietsman and Pistorius, 015) but often focus preferentially on heat transfer effects through the surrounding refractories since these are key factors in optimizing the engineering design of the tap-hole structure. Although it is a secondary problem in the thermal design, the fluid flow behaviour in furnace tap-hole entry regions and channels is nonetheless interesting and does impact operational aspects such as tap-hole refractory wear, tapping flow rates, and (in the case of single tap-hole furnaces) phase intermixing. As a result, further study of tap-holes from a fluid flow perspective alone was thought to be of some merit. Velocity and qualitative characteristics of the flow inside the tap-hole channel are functions of the furnace operating conditions, the geometry of the tap-hole, and the physical properties of the process materials. This was previously studied for the case of a generalized tap-hole carrying a single fluid phase (Reynolds and Erwee, 2017), and is extended with the present work to include the presence of porous media as well as multiphase fluid flow in the context of tapping from ferromanganese furnaces.
Fluid dynamics in furnace tap-holes is governed by the fundamental laws of conservation of momentum and mass, the Navier-Stokes and continuity equations. These nonlinear partial differential equations must be extended with terms to account for multiphase flow effects and drag forces that act on the flow field
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Software implementation
Solutions to Equations [1]–[4] were obtained using a finite volume method solver developed for the OpenFOAM® open source computational mechanics platform, version 5.0 (OpenFOAM, 2018). To account for phase dynamics the volume of fluid (VOF) method, in which a single velocity and pressure field describes both fluid phases present, was used. Velocity and pressure fields were solved at each time step using the PIMPLE algorithm, which unifies the PISO (pressure-implicit with splitting of operators) and SIMPLE (semi-implicit method for pressure-linked equations) approaches. Careful treatment of the porosity terms is necessary in order to retain numerical stability and accuracy near discontinuities, and is discussed at length in Oxtoby, Heyns, and Suliman (2013). Additional capability enabling the porosity field to follow only one of the fluid phases as the solution evolves in time was also implemented. Computational meshes for the models were generated using two different tools for the purposes of cross-comparison; ANSYS DesignModeler (ANSYS, 2018) and cfMesh (cfMesh, 2018). Models were executed in parallel on between 16 and 64 CPU cores using the CentOS Linux-based Lengau cluster at the Centre for High Performance Computing (CHPC 2018). Execution times ranged from two hours up to a day, depending on model resolution and parameters.
Model geometry and boundary conditions
Model description Governing equations
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Here the viscosities and densities refer to the fluid phase, and dc is the diameter of the particles making up the porous bed.
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In order to study the problem of fluid flow during ferromanganese tapping, a simple model geometry was developed consisting of a large plenum area representing the interior of the furnace and a cylindrical channel representing the tap-hole. Realistic dimensions typical of ferromanganese furnaces were used, and are reported later. The model geometry is shown schematically in Figure 3. The Journal of the Southern African Institute of Mining and Metallurgy
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Phase effects in tap-hole flow – a computational modelling study by the equivalent quiescent area. The concept is illustrated graphically in Figure 4. In addition to the quantitative information obtained, qualitative visualizations of the flow and phase fields were generated using the raw output from OpenFOAM® in ParaView, an open source visualisztion application (ParaView 2018).
Model parameter estimation
Figure 3—Geometry of model region showing boundary locations
Boundaries ‘A’ are treated as free inlet-outlet surfaces through which fluids can pass, and are taken to represent the farfield conditions in the furnace away from the tap-hole entrance area. P is fixed at the hydrostatic head values, u is calculated from the local pressure gradient, and a is set to fixed values representing the bulk level of the slag-metal interface far from the tap-hole. Boundaries ‘B’ are treated as solid walls, with a no-slip boundary condition applied for u and a fixed flux pressure (fixed hydrostatic gradient) condition for P. A zero-gradient boundary condition is also applied for the a field. Boundary ‘C’ is the taphole outlet at which P is set to atmospheric pressure, u is again calculated from the local pressure gradient, and a zero-gradient boundary condition is applied to a. All models were initialized at stagnant conditions with the level of the slag-metal interface equal to that in the bulk interior of the furnace.
Model post-processing In order to extract useful data outputs from the model, the results were processed at run time to obtain the values of slag mass . . flow rate ms, metal mass flow rate mm, and interface deformation δi as a function of time. The flow rates are obtained by simple integration on the computational mesh at the outlet, as shown in Equation [5].
Material properties and operating conditions are required as inputs to the computational models. In order to estimate material properties, their chemical compositions are required. A standard high-carbon ferromanganese product specification (Letsoalo and Steenkamp, 2017) was used for the metal, and a range of compositions from industrial operations as shown in Table I was used for the slag (Olsen, Tangstad, and Lindstad, 2007). Densities were estimated from molar volumes (Mills, 2011) and viscosities were calculated using structural methods available in the FactSage thermochemical software package (Bale et al., 2016; Tang and Tangstad, 2007). Coke particle sizes in the coke-slag bed in the vicinity of the tap-hole away from the central reaction zones under the electrodes were estimated from furnace dig-out photographs and previously published information to be under 20 mm (Barcza et al., 1979). The coke bed in this region is assumed to be very loosely packed, as it forms by gradual accumulation of coke particles at the periphery of the furnace by slow convection in the slag. For loose beds of irregular particles, packing porosities from 0.35 to 0.7 are possible (Yu, Zou, and Standish, 1996). In order to obtain an estimate of the internal hydrostatic pressure in the vicinity of the tap-hole, measures of the height of the burden layer, slag and coke bed, and metal phase above the tap-hole centreline were required. These were obtained from a combination of discussion with researchers working in the industry and previously published results (Barcza et al., 1979; Olsen, Tangstad, and Lindstad, 2007). The ranges used in the calculations are shown in Table II. Calculating the hydrostatic pressures from these values gives values between 50 and 90 kPa. Such a large range occurs primarily as a result of the significant variations in burden, slag
[5]
Here A is a vector directed normal to the surface of the mesh elements at the outlet boundary and with magnitude equal to their area, subscript n indicates values at a specific element in the surface mesh, and subscripts m and s refer to the metal and slag phases respectively. Calculation of the deformation of the interface is somewhat more difficult. At each time step, a small section of the mesh at the end of the tap-hole near to the outlet is first selected. The volume fraction of metal as well as the surface area of the slagmetal interface in this section is then calculated by integration in an analogous fashion to the method used for the flow rates. The equivalent surface area of the slag-metal interface under quiescent conditions is then calculated using the volume of the metal, and δi as a dimensionless measure of the deformation of the interface is obtained by dividing the actual surface area The Journal of the Southern African Institute of Mining and Metallurgy
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Figure 4—Diagram showing calculation of interface deformation, δi = A2/A1 for V1 = V2
Table I
Typical slag compositions for HCFeMn production Slag component
Mass % (min.)
SiO2 Al2O3 CaO MgO MnO VOLUME 119
Mass % (max.)
23.1 30.0 4.3 20.7 16.9 34.4 4.0 6.6 15.1 40.9 JUNE 2019
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Phase effects in tap-hole flow – a computational modelling study Table II
Furnace charge properties used to determine hydrostatic pressure for tapping Phase or layer in furnace Burden Slag Coke Bed Metal
Height 2.3–2.5 m 1.0–1.8 m 1.0–1.8 m 0–0.25 m
Density
Porosity
3000 kg/m3 0.4 3000 kg/m3 0.5 1500 kg/m3 0.5 7500 kg/m3 0
and coke, and metal heights that are possible between different furnaces and operating regimes. It is also important to note that this assumes that the burden layer is fully supported by the buoyancy of coke particles immersed in the slag (and possibly metal) layer, which is open to some debate.
resolution. The reason for this can be seen more clearly in Figure 6, which shows that the slag-metal interface breaks up considerably due to a highly three-dimensional swirling flow that develops as the metal enters the tap-hole underneath the more viscous, coke-filled slag layer. In VOF models, the minimum size of the entrained droplets of one phase in another is limited by size of the mesh elements; as the resolution is improved, more small droplets can be resolved and the calculated interface deformation increases. A summary of the mesh dependence results is shown in Figure 7.
Table III
ase case parameters for ferromanganese furnace B tap hole model Parameter
Results and discussion The operating conditions and material properties selected for the ferromanganese furnace tap-hole model are shown in Table III. The behaviour and numerical error of the model was initially studied for this base case set of parameters, followed by changes in individual values to identify sensitivity and operationalvariable responses. Here, ΔP is the hydrostatic head pressure applied as a boundary condition in the model, hi is the position of the interface in the main furnace volume relative to the tap-hole centreline, D is the tap-hole diameter, L is its length, and ht is the height of the centreline of the tapping channel above the furnace hearth. Each model was run for 3 seconds simulation time, with data being gathered and analysed over the last 0.5 seconds after the initial conditions had decayed and the flow was fully established. The coke bed was set to track the slag phase only, and was limited in extent to the plenum region outside the tap-hole channel.
μs ρs μm ρm γ f
Value
Parameter
Value
0.1 Pa.s 3000 kg/m3 0.005 Pa.s 7500 kg/m3 0.5 N/m 0.5
dc ΔP h i D L ht
0.01 m 70 kPa 0m 0.065 m 1.5 m 0.15 m
Numerical testing and validation comments Validation of computational models of single- and multiphase fluid flow in the presence of porous media has been addressed to a large degree in pre-existing work (Kadkhodabeigi, Tveit, and Johansen, 2011; Reynolds and Erwee, 2017; Oxtoby, Heyns, and Suliman, 2013), and the reader is referred to those publications for details. As the present work is primarily exploratory in nature, direct validation of the tap-hole model against measured data from ferromanganese furnaces is left to future work. It is, however, important to appreciate that complex nonlinear multiphysics models such as those used here can exhibit many undesirable dependencies on their numerical implementation; this is clarified for the ferromanganese tap-hole model in the following sections.
Figure 5—Evolution of slag and metal flow rates and interface deformation δi (see Figure 4) over time, comparison of mesh resolutions
Effect of model resolution In order to determine the mesh dependence of the model, meshes were generated at a number of different resolutions with characteristic element sizes in the tap-hole channel from 0.0025 m (239 066 elements) to 0.00125 m (1 841 291 elements). Results from the lowest and highest resolution cases are compared in Figures 5 and 6. It is clear that while the predicted flow rates of slag and metal are similar on both small and large meshes, the interface deformation shows considerable dependence on the model
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Figure 6—Slag-metal interface surface at end of simulation coloured by local fluid velocity, (a) low-resolution model, (b) high-resolution model The Journal of the Southern African Institute of Mining and Metallurgy
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Phase effects in tap-hole flow – a computational modelling study
Figure 7—Changes in calculated variables with mesh size (solid line shows average for last 500 ms of model data, shaded region shows standard deviation)
These results bear out the initial assessment that while the slag and metal flow rate behaviour is captured fairly consistently across a wide range of resolutions, the interface deformation continues to demonstrate appreciable mesh dependence up to the finest resolutions tested. This has implications for how such models are used – if the intention is to study interface effects in great detail, significantly larger meshes than those employed here will likely be necessary.
Effect of meshing algorithms The numerical characteristics of the model were investigated further using different software tools to generate the computational meshes used. Identical geometries were meshed using DesignModeler and cfMesh software at comparable linear element sizes and distributions, and exported for solution using the same OpenFOAM® solver in both cases. Although the meshes are constrained to produce elements of the same size in the plenum and tapping channel regions, the algorithms used in each tool differ to some degree and this results in slightly different mesh element arrangements and local refinements. This can be seen in Figure 8. Optimal meshing of geometries, particularly for hexahedral mesh elements, is tremendously complex and is still an open problem in the pre-processing of models for computational engineering; very few fully-automated methods exist for this task, hence the range of possible mesh topologies obtained from different tools. Models were run on meshes generated by both tools at tapping channel mesh element sizes of 0.0025 m and 0.00125 m. The results are shown in Figure 9. It is interesting to observe that at low resolutions, the time averages and dynamics in the results from both tools appear to match well for all output variables. However, at higher
Figure 8—Cross-sections of low-resolution meshes generated by (a) DesignModeler and (b) cfMesh, showing elements and refinement zones
resolutions, while the flow rates are correctly replicated between mesh types, it appears that the interface deformation is not. This suggests that on fine meshes not only is the absolute resolution important for determining the slag-metal interface distortion and breakup behaviour, but also the shape and local topology of the elements used in the mesh. The technique used to reconstruct the slag-metal interface surface and estimate its surface area also has a significant impact on the calculation of δi – the contour surface area method used here is known to be sensitive to mesh resolution and topology, and alternative methods more consistent with the underlying finite-volume formulation are currently being explored. Such aspects would need very careful consideration indeed should researchers wish to compare interface behaviour between different software platforms in the future.
Sensitivity analysis Although models at low resolution are not fully meshindependent and should therefore never be used for applications
Figure 9—Comparison of time evolution of model results using different mesh generation algorithms The Journal of the Southern African Institute of Mining and Metallurgy
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Phase effects in tap-hole flow – a computational modelling study requiring quantitative accuracy, their relative responses to changes in model parameters may still be expected to correctly indicate the sensitivities in the underlying problem physics. To achieve this for the ferromanganese tap-hole model, a large number of runs were performed using low-resolution DesignModeler meshes (tap-hole channel mesh element size 0.0025 m) keeping all parameters from Table III fixed except one, which was raised and lowered by 25% of its value. The response of the three outputs of the model – the time-averaged slag and metal flow rates and interface deformation over the last 500 ms run time – to each parameter change was then used to estimate a sensitivity relative to the base case model outputs. The complete sensitivity analysis was run for three major variations in the model – one per the default conditions in which the coke bed is located in the slag phase only, one in which the coke bed is allowed to extend into both slag and metal, and one in which the coke bed is entirely absent from the tap-hole entrance area.
The presence of the porous coke bed in the slag phase causes a partial obstruction to the material entering the tap-hole channel, and allows a higher flow rate of metal relative to slag to develop initially. This in turn causes the slag-metal interface to be forced upward as the metal enters beneath it, further restricting the slag flow. The end result is a metal-rich stream exiting the tap-hole, highly mixed with slag droplets due to the swirling flow and high shear rates that occur near to the entrance. Tapping mass flow rates for this case are 60.7 kg/s metal and 9.3 kg/s slag. The sensitivity of the model to material properties for this case is shown in Figure 11. The model exhibits very strong sensitivity to the coke bed parameters. Increasing the coke bed porosity or coke particle diameter results in an increase in both the slag flow rate and the interface deformation, and a smaller decrease in the metal flow rate. Of the remaining parameters the metal density and slag viscosity have a significant effect, while metal viscosity and interfacial surface tension have virtually no impact on the model behaviour.
Coke bed in slag phase only
Coke bed in slag and metal phases
For the default case in which the coke bed forms a porous medium in the slag phase only, the flow patterns inside the tap-hole channel are highly three-dimensional and complex. A selection of still frames from the model using base case parameters is shown in Figures 10a to 10c.
For the case in which the coke bed forms a porous medium around the tap-hole entrance in both the slag and metal phases, porous drag on the flow of both phases entering the tap-hole occurs. Still frames from this condition using base case model parameters are shown in Figures 12a to 12c.
Figure 10—Evolution of slag-metal interface over time, base case with coke in slag phase only
Figure 12—Evolution of slag-metal interface over time, base case with coke in slag and metal phases
Figure 11—Positive and negative sensitivities of model results to material properties, coke in slag phase only
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Phase effects in tap-hole flow – a computational modelling study It is immediately obvious when comparing to Figure 10 that the tapping velocities are significantly reduced when the coke bed obstructs both phases. While there is still some phase intermixing due to three-dimensional flow patterns near the channel entrance, this dissipates within a few tap-hole diameters and the slag and metal are well separated as they exit the tap-hole. Tapping mass flow rates for this case are 16.1 kg/s metal and 7.1 kg/s slag. The sensitivity of the model to material properties for this case is shown in Figure 13. As before, the strongest effects on the model behaviour are as a result of changes to the coke bed parameters. In this case increasing the porosity or coke particle diameter causes both the slag and metal flow rates to increase, and to a lesser degree the interface deformation. Metal density has a significant impact on metal flow rate, and slag density has a significant impact on slag flow rate. Slag viscosity affects both flow rates to a smaller degree, and the model is again insensitive to metal viscosity or interfacial surface tension.
No coke bed present For the case in which the coke bed is absent entirely from the tap-hole entrance region, only viscous and fluid acceleration forces limit the flow through the tap-hole. A series of still frames from this condition using base case model parameters is shown in Figures 14a to 14c. Comparisons with Figures 10 and 12 show that when coke is absent from the tap-hole entrance region, the velocities in the tapping flow are considerably higher. The increased velocities lead to the formation of Kelvin-Helmholtz shear instabilities at the slag-metal interface, which results in partial breakup and intermixing of the two fluids. Tapping mass flow rates for this case were 35.8 kg/s metal and 28.3 kg/s slag. The sensitivity of this model condition to material properties is shown in Figure 15. In this case, coke bed properties naturally do not affect the model. Of the remaining parameters, slag and metal densities cause the most significant changes in model outputs, affecting all three values. The fluid viscosities and interfacial surface tension have negligible effects on the flow rates, but are seen to change the interface deformation to a small degree.
Figure 14 – Evolution of slag-metal interface over time, base case with no coke present
Tap-hole wear effects While the base case model is representative of new ferromanganese tap-hole designs, tap-holes wear progressively throughout their lifetime and this can appreciably alter the geometry that the flow through the channel experiences at any given time. Tap-hole wear can vary greatly between designs, processes, and operating procedures (Nelson and Hundermark, 2016), and the way in which the geometry changes due to wear must therefore be assessed on a case-by-case basis. In manganese furnaces, dig-out evidence (Steenkamp et al., 2014) suggests that the mechanism of wear is discrete failure of successive tapblock inserts as a result of chemical and thermal attack at the hot face inside the furnace. This results in a progressive shortening of the tap-hole combined with a widening of the entire channel as a result of tap-hole opening operations.
Figure 13—Positive and negative sensitivities of model results to material properties, coke in slag and metal phases
Figure 15—Positive and negative sensitivities of model results to material properties, no coke bed present The Journal of the Southern African Institute of Mining and Metallurgy
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Phase effects in tap-hole flow – a computational modelling study The effect of this on the tapping flow behaviour was investigated by reducing L and increasing D in the low-resolution model geometry while keeping all other parameters in Table III fixed. The time-averaged results are shown in Figure 16. Mass flow rates of both slag and metal increase approximately quadratically with increasing tap-hole diameter, indicating that the linear velocity of the tapping stream remains reasonably constant. In contrast, tap-hole length has relatively little effect on the flow rates; although a small decrease would be expected as the tap-hole increases in length, the difference between the 0.5 and 0.75 m cases is so small that it is overwhelmed by statistical noise generated by the dynamic fluctuations in the flow. This is in agreement with results from single-phase tap-hole flow models, which indicate that tap-hole flows are generally dominated by fluid momentum rather than viscous effects (Reynolds and Erwee, 2017). The behaviour of the deformation of the slag-metal interface is more complex. The degree of interface deformation and mixing increases at larger diameters as the Reynolds number of the flow rises and the flow becomes more turbulent. There is also a strong decrease of δi with increasing tap-hole length, suggesting that most of the interface breakup occurs close to the tap-hole entrance – longer tap-holes therefore allow more time for slag and metal droplets to recombine and settle. Interface visualizations from worn tap-hole models are shown in Figure 17 and show the effect of shorter, wider tap-holes on the slag-metal interface. In cases of extreme wear and very large tap-hole diameters it is possible that resistances to fluid flow in the furnace interior may also start to play a role in determining the tapping flow rates and patterns. This would be particularly significant when there are obstructions to the flow (low-porosity packed bed structures, sintered or solidified process material, broken electrodes) in the vicinity of the tap-hole entrance. Such effects are not considered here, but should be regarded as factors that may influence tapping behaviour, particularly under abnormal operating conditions.
Operational parameters Changes in the hydrostatic pressure at the tap-hole inlet or interface position relative to the tap-hole centreline occur routinely during furnace operations. Pressures are affected by variations in charge composition and the heights of the slag, metal, and burden layers in the furnace. The slag-metal interface position typically changes over the duration of a single tap, with the interface starting above the tap-hole and dropping to level with, or even below it, as the furnace is emptied. In order to study the effect these operational parameters have on the behaviour of the model, cases covering a range of tapping pressures and interface positions were run while holding all
Figure 16—Changes in calculated variables with tap-hole length and diameter (solid line shows average for last 500 ms of model data, shaded region shows standard deviation)
other parameters at base case conditions, as per Table III. As the interfacial behaviour was of particular interest in these cases, the models used high-resolution computational meshes generated by DesignModeler (tap-hole channel element size 0.00125 m).
Tapping pressure ΔP was varied between 50 and 90 kPa, representing a likely operating range for a ferromanganese furnace. The effect of this change on the model outputs is shown in Figure 18. Flow rates increase approximately as the square root of the hydrostatic pressure applied at the plenum boundaries, which is in general agreement with the principle of momentum-limited fluid flow in the tap-hole (Nelson and Hundermark, 2016; Reynolds and Erwee, 2017). Together with this increase in flow rates and linear velocities, the interface deformation increases gradually as the local shear forces in the fluid grow larger and cause more disruption and breakup of the slag-metal interface. The irregular behaviour of δi is related to the large degree of variability in this measure in the model results, as evidenced by the large standard deviation.
Interface position between metal and slag Parameter hi was varied between 0.065 m above and below the tap-hole centerline (one tap-hole diameter in either direction). The response of the model to these changes is given in Figure 19.
Figure 17—Slag-metal interfaces in worn tap-holes, (l) L = 0.75 m, D = 0.065 m (r) L = 0.50 m D = 0.13 m
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Phase effects in tap-hole flow – a computational modelling study
Figure 18—Changes in calculated variables with applied pressure (solid line shows average for last 500 ms of model data, shaded region shows standard deviation)
Figure 19—Changes in calculated variables with initial interface location relative to tap-hole centreline (solid line shows average for last 500 ms of model data, shaded region shows standard deviation)
As the slag-metal interface inside the furnace moves downward past the tap-hole entrance, as it would over the course of a normal tap, the flow rate of metal through the tapping channel drops substantially while at the same time the flow rate of slag increases gradually. This net reduction in tapping flow rate occurs as a result of the porous coke bed in the slag layer obstructing flow into the tap-hole more and more as the interface drops past the tap-hole entrance. Visualizations of the interface shape and structure at the end of the simulation for different values of hi are shown in Figure 20. It is interesting to observe that the interface is drawn strongly toward the entrance as a result of the rapid flow acceleration into the tap-hole channel, even when it is well above or below the tap-hole centreline in the bulk interior of the furnace. This acceleration and change of direction causes a low-pressure region to develop at the tap-hole entrance, and this is sufficient to draw the interface into the channel from some distance above or below the tap-hole. Metal entrainment into the tap-hole is especially strong, and some metal may therefore be expected almost always to be drawn into the tapping stream. Interface deformation is also greatly increased as the slag-metal level in the furnace drops, predominantly as a result of the increasingly tortuous path the metal must take in order to enter the tap-hole. This causes additional shear stresses on the phase interface and results in more breakup and intermixing of slag and metal.
Conclusions A computational model of isothermal fluid flow through a ferromanganese furnace tap-hole was successfully developed and implemented using open source tools. Multiphase flow and porous media effects were included in the model to account for the presence of slag, metal, and solid coke particles in and around the tap-hole. The model’s numerical characteristics were tested and it was found that while fluid velocities and flow rates The Journal of the Southern African Institute of Mining and Metallurgy
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Figure 20—Slag-metal interfaces at end of simulation for different initial slag-metal levels
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Phase effects in tap-hole flow – a computational modelling study The model was used to study the sensitivity of the tapping problem to various parameters related to the physical properties of the slag and metal fluids as well as the coke bed. It was found that in cases where a coke bed is present, the porosity of the bed and coke particle size have the greatest impact on tapping flow rates and slag-metal interface distortion. The spatial distribution of the coke bed in the vicinity of the tap-hole was also seen to have a dominating effect on the flow patterns – when coke was present in the slag phase only, more three-dimensional swirling flow and interface breakup at the tap-hole entrance was observed compared to the cases in which the coke was either absent or extended into both slag and metal phases. The model was also applied to the effects of operational variables such as pressure and interface position variation during tapping, and tap-hole wear. Wider tapping channels typical of worn tap-holes were seen to significantly increase the flow rate of the tapping stream as well as the degree of slag-metal interface distortion, which is highest near to the tap-hole entrance. A greater hydrostatic pressure was observed to increase tapping flow rates in accordance with momentum-dominated rather than viscous-force-dominated fluid flow. The slag-metal interface in the furnace interior was also seen to be drawn down (or up) into the tapping channel from an appreciable distance above (or below) the tap-hole entrance; some mixing of slag and metal in the tapping stream is therefore likely to be inevitable throughout the duration of a tap. This study is by no means exhaustive, and is intended to be a starting point for more focused discussions and investigations. Further work in this area is suggested in the following areas: ➤ Problem-specific validation of models against data from ferromanganese tapping operations ➤ Refinement of the understanding of the nature and location of the coke bed in the vicinity of the tap-hole ➤ Refinement of worn tap-hole geometries, including the possibility of a priori prediction using computational methods ➤ More extensive cross-verification of modelling results between meshing and software packages to determine appropriate settings for reliable comparisons ➤ Exploration of numerical interface compression schemes, interfacial mesh refinement algorithms, and sub-grid scale models to reach full mesh independence for reasonable computational cost ➤ Assessment of the risk/reward of including additional physics such as heat transfer, phase change, and thermochemical effects in computational models.
Spencer, P. , and van Ende, M.-A. 2016. FactSage thermochemical software and databases 2010–2016. Calphad, vol. 54. pp. 35–53. Barcza, N.A., Koursaris, A., See, J.B., and Gericke, W.A. 1979. The ‘dig out’ of a 75 MVA high-carbon ferromanganese electric smelting furnace. Proceedings of the 37th Electric Furnace Conference, Detroit, MI, 4–7 December 1979. American Institute of Mining, Metallurgical, and Petroleum Engineers, New York. pp. 19-33. cfMesh. 2018. Creative Fields. https://cfmesh.com/ CHPC. 2018. Centre for High Performance Computing Lengau Cluster. https://www. chpc.ac.za/index.php/resources/lengau-cluster Degel. R., Fröhling. C., Köneke. M., Hecker. E., Oterdoom. H., and van Niekerk. A. 2015. History and new milestones in submerged arc furnace technology for ferro alloy and silicon production. Proceedings of the Fourteenth International Ferro-Alloys Congress, INFACON XIV, Kyiv, Ukraine, 1–4 June 2015. pp. 7–16. https://www.pyrometallurgy.co.za/InfaconXIV/007-Degel.pdf Kadkhodabeigi. M., Tveit, H., and Johansen, S.T. 2011. Modeling the tapping of silicon melt from the submerged arc furnaces. ISIJ International, vol. 51, no. 2. pp. 193–202. Letsoalo, M. and Steenkamp, J. 2017. Round robin standard composition: HCFeMn reference material. Proceedings of the 3rd SAIMM Young Professionals Conference, Pretoria, South Africa, 9–10 March 2017. Southern African Institute of Mining and Metallurgy, Johannesburg. pp. 355–357. Mills, K.C. 2011. The estimation of slag properties. Short course given as part of SAIMM Pyrometallurgy 2011, Cradle of Humankind, South Africa, 6-9 March 2011. http://www.pyrometallurgy.co.za/KenMills/ Muller, J., Zietsman, J.H., and Pistorius, P.C. 2015. Modelling of manganese ferroalloy slag properties and flow during tapping. Metallurgical and Materials Transaction B, vol. 46B. pp. 2639–2651. Nelson, L.R. and Hundermark, R.J. 2016. The tap-hole – key to furnace performance. Journal of the Southern African Institute of Mining and Metallurgy, vol. 116, no. 5. pp. 465–490. Olsen. S.E., Tangstad. M., and Lindstad T. 2007. Production of Manganese Ferroalloys. Tapir Academic Press, Trondheim, Norway. OpenFOAM. 2018. The OpenFOAM Foundation.https://openfoam.org/ Oxtoby, O.F., Heyns, J.A., and Suliman, R. 2013. A finite-volume solver for two-fluid flow in heterogeneous porous media based on OpenFOAM® . Proceedings of the Open Source CFD International Conference 2013, Hamburg, Germany, 24–25 October 2013. pp. 1–12. ParaView. 2018. ParaView. https://www.paraview.org/ Reynolds, Q.G. and Erwee, M.W. 2017. Multiphase fluid flow modelling of furnace tap-holes. Proceedings of the 12th International Conference on CFD in Oil & Gas, Metallurgical and Process Industries, Trondheim, Norway, 30 May – 1 June 2017. Sintef, Trondheim. pp. 521–530. http://www.opensourcecfd.com/ Steenkamp, J.D., Gous, J.P., Pistorius, P.C., Tangstad, M., and Zietsman, J.H. 2014. Wear
This paper is published by permission of Mintek. The assistance of our colleagues Dr Joalet Steenkamp, Dr Stein Tore Johansen, and others in discussing the problem statement and determining typical parameters for ferromanganese furnaces is gratefully acknowledged. Resources provided by the CSIR/Meraka Institute Centre for High Performance Computing were invaluable in the completion of this study.
analysis of a tap-hole from a SiMn production furnace. Proceedings of the SAIMM Furnace Tapping Conference, Muldersdrift, South Africa, 27–29 May 2014. Southern African Institute of Mining and Metallurgy, Johannesburg. pp. 51–62. Tang, K. and Tangstad, M. 2007. Modelling viscosities of ferromanganese slags. Proceedings of the Eleventh International Ferro-Alloys Congress INFACON XI, New Dehli, India, 18–21 February 2007. Indian Ferro Alloy Producers Association. pp. 344–357.
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Acknowledgements
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Wear analysis of tap-holes at two ferrochromium production furnaces
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J.D. Steenkamp12
Affiliation: 1 Mintek, Randburg, South Africa. 2 University of the Witwatersrand, South Africa.
Synopsis
joalets@mintek.co.za
In July of 2016 and June 2017, major tap-hole repairs were carried out during the annual shutdown of two 63 MVA submerged arc furnaces producing ferrochromium in South Africa. The hot excavations of the tap-holes allowed for a study of the wear profiles. Subsequent thermodynamic calculations allowed for the quantification of the potential for chemical reaction between refractory and slag or alloy contributing to wear in the tap-hole area. Results from both wear profiling and thermodynamic calculations are reported here.
Dates:
Keywords ferrochromium furnace, tap-hole, wear analysis, refractory corrosion.
Correspondence to: J.D. Steenkamp
Email:
Received: 14 Mar. 2019 Revised: 16 Apr. 2019 Accepted: 6 May 2019 Published: June 2019
How to cite:
J.D. Steenkamp. Wear analysis of tap-holes at two ferrochromium production furnaces. The Southern African Insitute of Mining and Metallurgy DOI ID: http://dx.doi.org/10.17159/24119717/669/2019 ORCiD ID: J.D. Steenkamp https://orchid.org/0000-00030635-7927
Introduction Ferrochrome (FeCr) is an essential alloy in the production of stainless steel (Gasik, 2013). In 2017, 11.7 Mt of FeCr was produced (Pariser et al., 2018.). The main producers were China (39%), South Africa (29%), Kazakstan (12%), and India (8%). A number of grades of FeCr are available to producers of stainless steel, including high-carbon ferrochrome (HCFeCr) (Gasik, 2013). HCFeCr typically contains 60–70% Cr and 4–6% C (Basson and Daavittila, 2013). When the chromium content is lower (50–55% Cr) and the carbon content higher (6–8% C), it is referred to as charge chrome. Charge chrome is typically produced in South Africa due to the grades of ore available. FeCr is produced in submerged arc furnaces (SAFs) and in open arc furnaces (Mc Dougall, 2013) by carbothermic reduction of oxide raw materials. In a SAF the electrode tips are submerged in a porous charge mix. In FeCr production, the electrical energy is mainly liberated by resistive heating of the wet coke bed (Barker et al., 2007; Steenkamp et al., 2017). Thermally conductive lining design philosophies are typically applied in SAFs producing FeCr (Steenkamp, Denton, and Hayman, 2017; Coetzee, Duncanson, and Sylven, 2010; Coetzee et al., 2010). In July of 2016 and June of 2017, major tap-hole repairs were carried out during the annual shutdown of two 63 MVA SAFs producing FeCr in South Africa. The hot excavations of the tap-holes allowed for a study of the wear profiles. Both SAFs were operated on the same site, by two different operating crews, using the same raw materials. The major difference between them was that the SAF excavated in 2016 was an open SAF (referred to as Furnace A) and the one excavated in 2017 was a closed SAF (referred to as Furnace B). In an open SAF, a gap exists between the top of the steel shell and the roof, with the top of the burden being open to the atmosphere and the CO-rich off-gas being combusted as it escapes. In a closed SAF, the roof is attached to the steel shell and the CO-rich off-gas is collected from an offtake on the roof for further treatment and used in downstream processes. The paper presented here reports the macro-scale observations on refractory wear for the two SAFs, and studies the potential for chemical wear of refractory by slag and alloy as a potential contributing factor to the tap-hole wear observed.
Background Refractory wear mechanisms associated with the tap-hole region The refractory wear mechanisms reported in the literature for SAFs are densification, spalling, erosion, and corrosion. The Journal of the Southern African Institute of Mining and Metallurgy
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➤
Wear analysis of tap-holes at two ferrochromium production furnaces Densification (infiltration) is caused by slag or metal infiltrating pores and/or reacting with refractory (Hancock, 2006; Inada et al., 2009). Spalling is caused by thermal stress due to a high thermal gradient across a single refractory body (Hancock, 2006). Refractory material on the hot face fractures and breaks away due to densification and/or thermal stress (Coetzee, Duncanson, and Sylven, 2010). Erosion (abrasion) is caused by slag, metal, and solid material abrading refractory (Hancock, 2006; Duncanson and Toth, 2004). Corrosion (chemical wear) is caused when the system is not at equilibrium. Slag, metal, or a flux will dissolve the refractory components it is not saturated with and/or chemically react with them (Hancock, 2006). Corrosion always begins with the interaction of some minor (binder phase) or major (aggregate) constituents of the refractories with the liquid slag or metal (Hubble, 1999). Similarly, gas reactions between carbon refractory materials and CO2, H2O, and O2 (pure and in air) have been reported, as summarized by Nelson and Hundermark (2016). Tap-hole wear is a complex phenomenon (Nelson and Hundermark, 2016) and although oxygen lancing is commonly seen as the main contributor to tap-hole wear (Erwee, Reynolds, and Zietsman, 2016), studying the potential for other wear mechanisms (i.e. the potential for corrosion of tap-hole refractory by slag or alloy during tapping) assists in the selection of refractory materials used to construct tapblocks to minimize the potential for wear.
Figure 1—Sketch in plan view, indicating the conceptual layout of the two SAFs under investigation. The tap-hole repair area is indicated by the dashed red rectangle
Refractory design The SAFs under investigation were of circular design, with inner diameters of the steel shells at 13.7 m and heights (from the cold face of the hearth refractories to the top of the sidewalls) approximately 5.5 m. The three Søderberg electrodes (1.55 m in diameter) were positioned equilaterally with one, single-level taphole positioned at electrode 1 (referred to as the ‘tap electrode’) – see Figure 1. For both shutdowns, preparations were made for 3 × 3 m repair of the refractory material in the tap-hole area. In the case of Furnace A, the sidewall at clover no. 2 was also repaired, and in the case of Furnace B, the sidewalls at both clover no. 2 and clover no. 3 were repaired. The refractory designs in the tapblock areas are indicated in Figure 2. In both designs, a single carbon block formed the tapblock on the inside of the steel shell. On the cold face of the tapblock a smaller replaceable block (the quick-replaceable carbon block in Figure 2) was installed. In the case of Furnace B, the tapblock was extended further into the furnace by the installation of a row of carbon bricks towards the hot face of the tapblock. The hot face of the carbon bricks was protected by a layer of sacrificial alumina bricks. Towards the cold face, semi-graphite bricks and graphite bricks or tiles were installed. Semi-graphite bricks have artificial graphite as aggregate and petroleum pitch or coal tar as binder. No copper coolers were installed in the tap-holes, and the steel shells around the tapholes were cooled with water – thin-film cooling in the case of Furnace A and forced channel cooling in the case of Furnace B.
Tapping practice With the furnaces having only one single-level tap-hole, slag and alloy were tapped simultaneously. Each furnace was equipped with a pedestal-type drill-and-claygun. Furnace A was tapped
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Figure 2—FeCr lining design in the tap-hole area for (a) Furnace A and (b) Furnace B – drawings to scale
eight times a day on average with a standard deviation of 1 (calculated from a daily production data-set for 1 January 2014 to 30 June 2016). Similar information for Furnace B was not available. Operational practices included the sampling of slag and alloy at each tap. The slag sample was taken with an oxygen lance in the launder because oxygen lances were available freely on the tap floor. The alloy sample was taken in the alloy ladle with a ‘lollipop–sample dipstick’ or also in the launder. Slag and alloy compositions were determined by powder X-ray fluorescence (XRF). The carbon content of alloy samples was determined by combustion (LECO). The calculated averages and standard deviations in chemical compositions of tapped slag (six main components only) and alloy (four main components only) were calculated for the periods The Journal of the Southern African Institute of Mining and Metallurgy
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Wear analysis of tap-holes at two ferrochromium production furnaces 1 January 2014 to 30 June 2016 for Furnace A and I January 2012 to 5 July 2017 for Furnace B. The data-sets were not filtered for outliers. The alloy compositions are reported in Table I, and slag compositions in Table II. Tap temperatures were not measured on either of the SAFs. Liquidus temperatures were calculated by the plant operating system based on the chemical compositions of the slag and alloy. The calculations used to determine these liquidus temperatures were not made available to the author. The calculated averages and standard deviations in liquidus temperatures of the slag and alloy are summarized in Table III. Data-sets for the same periods were used and were again not filtered for outliers.
Tapblock life The tapblock life of Furnace B was significantly shorter than for Furnace A – see timelines in Figure 3. During initial installation, the SAFs were fully relined. During a tapblock repair only the tapblock is replaced, while during a partial rebuild both the tapblock and a section of the sidewall are replaced. The linings were installed in 2005 in Furnace A and 2011 in Furnace B. In Furnace A, the tapblock was replaced during a repair on average every 2–3 years. In Furnace B, the tapblock was replaced either during a repair or during a partial re-build, on average every 18 months.
Method Furnace preparation prior to switch-out In both instances, when the SAFs were prepared for the partial rebuild, the intention was for the SAFs to be emptied by melting down the burden and tapping through the tap-hole. For a specific SAF, the night prior to switch-out, feed to the furnace was stopped and power input maintained. The melt-out was monitored by managing the furnace power input (MW) and specific energy consumption (MWh per ton alloy tapped). During such an operation, not only the burden (consisting of ore, flux, reductant, slag, coke bed, and alloy) will be melted and tapped, but also any freeze lining that formed during normal operation of the furnace.
Tap-hole wear study on the macro-scale and wear profiling For the macro-scale investigation, photographs were taken of the refractory in situ. A CANON EOS 30D camera and CANON zoom lens (EF 75–300 mm F4–5.6) installed on a tripod and triggered by remote trigger was used. The camera settings for aperture and shutter speed were adjusted manually based on the lightmeter readings on the camera. Lighting was provided by free-standing floodlights and no flash was applied. The refractory thickness was measured using a tape measure.
Table I
Calculated average and standard deviations of the four-component alloy compositions (mass per cent) per furnace. For Furnace A, calculations are based on 7362 data points (1 January 2014 to 30 June 2016). For Furnace B, calculations are based on 13 686 data points (1 January 2012 to 5 July 2017) Average Standard deviation
alculated average and standard deviations of the C calculated alloy and slag liquidus temperatures (°C). For Furnace A, calculations are based on 7362 data points (I January 2014 to 30 June). For Furnace B, calculations are based on 13 686 data points (I January 2012 to 5 July 2017)
Furnace A
Cr Si Fe C Total 49.9 0.9
Cr Average Standard deviation
Table III
4.8 0.9
Furnace B
50.1 0.6
36.7 1.0
Si
Fe
4.5 1.1
36.4 0.8
6.5 0.3
C
98.0 -
Total
6.7 0.5
Alloy
Slag
Average StDev Average StDev
Furnace A Furnace B
1550 1551
134 12
1654 1668
31 43
97.8 -
Table II
Calculated average and standard deviations of the six-component slag compositions (mass per cent) per furnace. For Furnace A, calculations are based on 7477 data points (1 January 2014 to 30 June 2016). For Furnace B, calculations are based on 14 082 data points (1 January 2012 to 5 July 2017) Furnace A
Al2O3 CaO Cr2O3 FeO MgO SiO2 Total Average Standard deviation
27.5 1.4
4.2 0.7
12.1 2.0
5.7 1.0
18.1 1.0
30.5 1.9
98.0 -
Furnace B
Al2O3 CaO Cr2O3 FeO MgO SiO2 Total Average Standard deviation
27.6 1.8
5.4 1.7
10.9 2.0
5.2 1.1
18.9 1.2
30.0 2.6
98.0 -
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Figure 3—Timelines indicating tapblock life and repair activities related to tapblock life for (a) Furnace A and (b) Furnace B. Questions marks indicate activities for which dates were estimated by operating crews as no records were available VOLUME 119
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Wear analysis of tap-holes at two ferrochromium production furnaces Thermodynamic calculations approach The containment philosophy for the two furnaces was based on a conductive lining design (Steenkamp, Denton, and Hayman, 2017). In a conductive lining design where a layer of process material is frozen on the hot face of the refractory lining, chemical compatibility between refractory material and liquid process material is not important as the intention is for a frozen layer of process material to form at the interface between the liquid process material and the refractory material. Due to the dynamic conditions in the tap-hole area, the probability of a stable freeze lining forming is very low, as the excavation of these two tapblocks clearly illustrates. The probability that during tapping the tapblock refractory material is in contact with liquid alloy or slag is very high. Therefore, the potential for chemical wear of the tapblock refractory by FeCr alloy or slag was determined by thermodynamic calculation. The Equilibrium model in FACTSage 7.2 was applied in the thermodynamic calculations (Bale et al., 2002). Depending on the type of calculation the FToxid and/or SGTE and FACTPS databases were selected. Default gas and solids were selected as pure species. Duplicates were suppressed with the order of preference being FACTPS, FToxid, and SGTE databases. The composition of the alloy used in the calculations is presented in Table IV, and slag in Table V. In all calculations, initial conditions were not specified. The alloy composition in Table IV is the average of the normalized alloy compositions for Furnace A and Furnace B presented in Table I. The slag compositions in Table V were determined as follows: ➤ Slag 1 is the average of the normalized slag compositions for Furnace A and Furnace B presented in Table II. These analyses showed elevated levels of Cr2O3 and FeO due to the presence of entrained alloy as well as partially altered chromite (Bergmann, Govender, and Corfield, 2016). Using the results in thermodynamic calculations will result in levels of wear that are artificially attributed to the reduction of FeO, and to a lesser extent reduction of Cr2O3. ➤ Slag 2 is Slag 1 without the Cr2O3 and FeO present and normalized. The purpose was to determine if reduction of any of the other phases present in the slag could result in refractory wear. ➤ Slag 3 is Slag 1 excluding the presence of alloy phases, based on a study done by Bergmann, Govender, and Corfield (2016). For the purpose of the calculations, the composition of the carbon block was assumed to be 100% C as the actual composition of the refractory was not known. As an alternative, refractory material consisting of 100% SiC was considered due to the use of SiC-based refractory material as a tapblock in silicomanganese production (Steenkamp, Pistorius, and Muller, 2016).
The temperature range was selected based on the calculated liquidus temperatures of the tapped alloy and tapped slag presented in Table III. For alloy, the temperature range 1450– 1650°C at 50°C intervals was selected, and for slag, 1550–1750°C at 50°C intervals. In all instances, results were calculated at the temperatures selected as well as at temperatures where phase transformations occurred, i.e. normal+transitions was selected. A pressure of 1 atmosphere was selected, although the ambient pressure at the plant is typically 0.85 atmospheric.
Equilibrium phase distributions of alloy and slag on their own To obtain an understanding of the system under investigation, the equilibrium phase distributions of slag and alloy were calculated. For slag the FToxid and FACTPS, and for the alloy the SGTE and FACTPS databases were selected. For alloy, all solution phases were selected except for SGTE-DIAM. For slag, all solution phases were selected except for FToxid-OlivB and FToxid-Oliv? (one of the solution phases available in the FToxid database).
Equilibrium phase distribution of alloy reacted with refractory The equilibrium phase distribution of the reaction products for the reaction of 50 g of alloy with 50 g of refractory was then calculated as a function of temperature. The aim was to identify the potential for chemical reactions between refractory and alloy, the types of reactions that would occur, and the temperatures at which these reactions would occur. The SGTE and FACTPS databases were selected. All solution phases were selected except for SGTE-DIAM. For pure solids, Cr4C formation was suppressed to allow for the formation of carbide solution phases.
Equilibrium phase distribution of slag reacted with refractory Subsequently, the equilibrium phase distribution of the reaction products for the reaction of 50 g of slag with 50 g of refractory was calculated as a function of temperature. The aim was to identify the potential for chemical reaction between refractory and slag, the types of reactions that would occur, and the temperatures at which these reactions would occur. The SGTE, FTOxid, and FACTPS databases were selected. All solution species were selected except for FToxid-OlivB, FToxid-oliv?, and SGTE-DIAM. For pure solids Cr4C formation was suppressed.
Refractory consumption Finally, refractory consumption as a function of temperature was calculated according to Equation [1]. [1]
Table V
omposition (expressed in mass per cent) of slag C applied in free energy minimization calculations
Table IV
Composition (expressed in mass per cent) of alloy applied in free energy minimization calculations Alloy 1
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Slag 1 Slag 2 Slag 3
28.1 34.3 31.4
CaO 4.3 5.2 4.8
Cr2O3 FeO MgO SiO2 Total 12.3 0.0 9.4
5.8 0.0 1.7
18.4 22.5 20.7
31.1 38.0 32.0
100.0 100.0 100.0
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Wear analysis of tap-holes at two ferrochromium production furnaces where: CR: Refractory consumption, in mass per cent MO: Original mass of refractory, in grams ME: Mass of refractory in equilibrium with alloy or slag at a specific temperature, in grams
Results and discussion Plant-based observations The tap-hole area of Furnace A was in a much better condition than that in Furnace B, as seen in Figure 4a compared to Figure 4b. The photographs in Figure 4 were taken after removal of the tap launders. At both SAFs, the quick-replaceable carbon blocks were installed in an area referred to as the chapel, which was named after the shape of the protective heat shield around this area. In Figure 4, parts of the steel shells of the protective heat shields are clearly visible as the refractory materials which cover them during operation were removed by the time the photographs were taken. In Figure 4a the rubble generated during the removal of the launder and the refractory from the protective heat shields is still visible. The quick-replaceable carbon blocks on both SAFs had very large, oval-shaped tap-holes as seen in Figure 4. The tap-hole in Furnace A was filled with tap-hole clay (Figure 4a) and in Furnace B with slag and alloy (Figure 4b). For each SAF, the oval-shaped tap-hole appeared to consist of two distinct tap-holes, as is clearly visible in Figure 4a. Below the quickreplaceable carbon block a third ‘tap-hole’ was visible, again seen more clearly in Figure 4a. These observations indicate that even though the tap-hole design philosophy was based on one
Figure 4—Overview of tap-hole face after removal of the launder and refractory material covering the chapel for (a) Furnace A (view in photograph is perpendicular to the tap-hole face) and (b) Furnace B (view in photograph is around 45° offset to the right of the line perpendicular to the tap-hole face) The Journal of the Southern African Institute of Mining and Metallurgy
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single-level tap-hole, normal operations require two tap-hole levels and emergency operations a third. Therefore the tap-hole life-cycle (Steenkamp et al., 2016) was not taken into account when the refractory design was undertaken. Furnace designers will probably argue that operations were not conducted properly, but these two SAFs were operated by two different operating crews and the fact that the observation was made on both SAFs is probably an indication that tapblock life could be improved by a design that allows for the need for three tap-hole levels. The dimensions of the quick-replaceable carbon blocks were determined by measuring tape, and those of the oval-shaped taphole by dimensional analysis (see Figure 5). In both cases, the oval-shaped tap-holes covered in the order of 50% of the surface area of the quick-replaceable carbon blocks. In comparison, for a new quick-replaceable carbon block the tap-hole area is only 10% of the total area. Another observation made from Figure 5 is that the quickreplaceable carbon block in Furnace B was installed at a 90° offset from the original design. In Figure 5a the long end of the block is in a horizontal position and in Figure 5b in a vertical position. The deviation is attributed to the tapblock wear in Furnace B being much more significant than in Furnace A, as can be seen in Figure 6. Upon removal of the quick-replaceable carbon blocks, very little of the original tapblock remained for Furnace A (Figure 6a) and none for Furnace B (Figure 6b). The areas where the tapblocks were supposed to be were filled with process material. In Furnace A (Figure 6a), the process material consisted mainly of coke bed, in other words slag and carbonaceous reductant. In Furnace B (Figure 6b), the process material consisted of a coke bed and tap-hole clay mixture towards the top and alloy at the bottom. In Figure 6b, the tap-hole clay used to plug the tap-hole is clearly visible. For Furnace A, a wear profile of the refractory surrounding the tapblock could be determined by manual measurement once the refractory around the tapblock cooled, allowing access. Measurements were done using a tape measure and the results
Figure 5—Dimensions of the quick-replaceable carbon blocks and large oval tap-holes for (a) Furnace A and (b) Furnace B VOLUME 119
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Wear analysis of tap-holes at two ferrochromium production furnaces Equilibrium phase distribution of alloy reacted with refractory In Figure 9, the results for alloy in contact with refractory are presented. For alloy in contact with C-based refractory, three phases are present: the C-based refractory, the liquid alloy, and an ironcontaining chrome carbide with an increased carbon content (metal-cation-to-carbon ratio is 1.5 for the carbide formed in Figure 9a, compared to 2.3 for the carbide formed in Figure 8a). This carbide phase is present only at temperatures below 1475°C. The slight reduction in the amount of C-based refractory is either due to the formation of (Cr, Fe)3C2 or to solution into the liquid alloy.
Figure 6—Inspection just after removal of the quick-replaceable carbon blocks revealed that very little of the tapblock remained in (a) Furnace A and (b) Furnace B
Figure 7—Numbered courses of bricks to the right-hand side of the tap-hole with refractory wear profile (red line) superimposed onto refractory design drawing
were superimposed onto the design drawing – see Figure 7. As can be seen from the drawing, only in the order of 50% of the sidewall refractory surrounding the tapblock remained. The wear observed was probably a result of a combination of sidewall refractory wear associated with each of the clovers and wear associated with the tapping of liquid slag and alloy through the tap-hole and associated lancing activities. For Furnace B, a wear profile could unfortunately not be determined due to time constraints.
Thermodynamic calculations In the sections below, the phase distributions for specific conditions are presented as a function of temperature. The conditions were alloy or slag on their own, and slag or alloy in combination with C or SiC refractory.
Figure 8—Equilibrium phase distribution as a function of temperature for only (a) Alloy 1 or (b) Slag 1 to Slag 3
Equilibrium phase distributions of alloy and slag on their own In Figure 8, the results for alloy and slag only are presented. At temperatures below 1500°C a solid iron-containing chrome carbide is present (see Figure 8a). Above 1500°C the alloy will be 100% liquid. In the slags, a spinel (Cr, Fe, Mg, and Al-rich) and a liquid slag phase are present in the temperature range 1550°C to 1750°C. As the temperature increases, the liquid slag phase increases as expected. The slag:spinel ratios for Slag 1 and Slag 3 are very similar over the temperature range, ranging between 2.3 and 4.5. For Slag 2, the ratio is higher over the temperature range and ranges between 3.9 and 7.8. The lower spinel content in this slag is expected due to the absence of spinel-forming Cr and Fe.
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Figure 9—Equilibrium phase distribution as a function of temperature for Alloy 1 and (a) C-based refractory and (b) SiC-based refractory at an alloy:refractory mass ratio of 1:1 The Journal of the Southern African Institute of Mining and Metallurgy
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Wear analysis of tap-holes at two ferrochromium production furnaces For alloy in contact with SiC-based refractory, only two phases are present: the SiC-based refractory and the liquid alloy. In the temperature range under investigation, the reduction in the amount of SiC-based refractory is due to solution into the liquid alloy. From these results, it can be concluded that two potential corrosion mechanisms apply: chemical reaction or dissolution, described by the following three chemical reactions: [2] [3] [4] * C can also be in solution as per Equation [2]
where: <M> depicts species present as a solid phase (M) depicts species present as a liquid phase M depicts species in solution M depicts species present as a gas phase.
The liquid slag phase is present in all three cases in the temperature range under investigation. Only for Slag 2 and Slag 3 are there spinel slag phases, and only at temperatures below 1600–1650°C. For Slag 1, C forms as a reaction product below 1580°C, and for Slag 3 SiC forms below 1600°C. In the slags with Cr and Fe present (Slag 1 and Slag 3), liquid alloy forms in the temperature range under investigation. Gas also forms in increasing amounts as the temperature increases. For Slag 2 no gas forms as no Cr2O3 or FeO are present in the slag. From these results, it can be concluded that one potential corrosion mechanism applies: chemical reaction, described by Equations [3] and [4] above, and Equations [8] and [9]: [8] [9] ** Dissolution of the spinel phase will play a role, not indicated here *** SiC can also be in solution as per Equation [3]
Refractory consumption
Equilibrium phase distribution of slag reacted with refractory In Figure 10, the results for slag in contact with C-based refractory are presented. For all three slags, spinel and liquid slag phases are present in the temperature range under investigation. Interestingly enough, the slag:spinel ratio initially increases and then decreases with temperature in all three instances. For all three slags, SiC forms as a reaction product at around 1600–1650°C. The amount of SiC formed is greatest for Slag 2 (see Figure 10b). In the slags with Cr and Fe present (Slag 1 and Slag 3), liquid alloy forms in the temperature range under investigation. For Slag 3, the iron-containing chrome carbide with metalcation-to-carbon ratio of 2.3 observed in Figure 9a also forms at temperatures below 1590°C. From these results, it can be concluded that one potential corrosion mechanism applies: a chemical reaction – described by Equations [5] to [7]:
From the results presented in Figure 9, the main corrosion mechanism responsible for the wear of C- or SiC-based refractory by FeCr alloy is the dissolution of the refractory into the alloy. As can be seen in Figure 12, the potential for SiC-based refractory to dissolve into alloy is greater than for C-based refractory. From results presented in Figure 10 and Figure 11, the main corrosion mechanism responsible for the wear of C- or SiCbased refractory by FeCr slag is chemical reaction. This includes the reduction of CrO and FeO in the slag to form alloy, and the reduction of SiO2 in the slag to SiC.
[5] [6] [7] ** Dissolution of the spinel phase will play a role, not indicated here
Figure 11—Equilibrium phase distribution as a function of temperature for SiC-based refractory: (a) Slag 1, (b) Slag 2, and (c) Slag 3
In Figure 11, the results for slag in contact with SiC-based refractory are presented.
Figure 10—Equilibrium phase distribution as a function of temperature for C-based refractory: (a) Slag 1, (b) Slag 2, and (c) Slag 3 The Journal of the Southern African Institute of Mining and Metallurgy
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Figure 12—Refractory consumption by alloy for (a) C-based refractory and (b) SiC-based refractory VOLUME 119
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Wear analysis of tap-holes at two ferrochromium production furnaces For Slag 2, where CrO and FeO are absent, only SiC formation applies and only for C-based refractory material. For Slag 1 and Slag 3, both alloy and SiC formation apply. In alloy formation, dissolution of the spinel phase into the slag to supplement the CrO and FeO, reduced according to Equations [6] to [9], plays a role. To what extent the reaction kinetics will allow for the dissolution to occur during tapping is a matter for further investigation. Based on the refractory consumption calculations presented in Figure 13, C-based refractory as a tapblock material will perform similarly to SiC-based refractory at temperatures below 1600°C. At 1600°C to 1650°C, SiC formation will result in increased consumption of C-based refractory, and at temperatures exceeding 1650°C the SiC-based refractory will outperform the C-based refractory. To what extent lancing affects the temperature and slag composition in the tapblock, and how this contributes to the mechanisms observed, will be interesting to understand.
➤ Further investigation into the effect of the lancing process on tap-hole wear would add significant value to the understanding of tap-hole wear, both in terms of the effect of the reaction between the refractory and oxygen in the gas phase and also in terms of the changes in chemical composition of the slag and the temperature changes in and around the tapblock.
Conclusions
Bale, C., Chartrand, P., Degterov, S., and Eriksson, G. 2002, FactSage thermochemical software and databases, Calphad, vol. 26. pp. 189–228. Barker, I.J., Rennie, M.S., Hockaday, C.J., and Brereton-Stiles, P.J. 2007. Measurement and control of arcing in a submerged-arc furnace. Proceedings of Infacon XI: Innovation in Ferroalloy Industry, New Delhi, India, 18-21 February 2007. Indian Ferro Alloy Producers Association. https://www.pyrometallurgy.co.za/ InfaconXI/685-Barker.pdf Basson, J. and Daavittila, J. 2013. High carbon ferrochrome technology. Handbook of Ferroalloys – Theory and Technology. Butterworth-Heinemann, Oxford, UK. Chapter 9. Bergmann, C., Govender, V., and Corfield, A.A. 2016. Using mineralogical characterisation and process modelling to simulate the gravity recovery of ferrochrome fines. Minerals Engineering, vol. 91. pp. 2–15. Coetzee, C., Duncanson, P.L., and Sylven, P. 2010. Campaign extensions for ferroalloy furnaces with improved tap hole repair system. Proceedings of Infacon XII: Sustainable Future, Helsinki, 6-9 June 2010. Outotec Oyj. pp. 857–866. Coetzee, C., Lamont, P.H., Duncanson, P.L., and Sylven, P. 2010. New refractory lining direction at Jindal Stainless FeCr#1 and #2 furnaces. Proceedings of Infacon XII: Sustainable Future, Helsinki, 6–9 June 2010. Outotec Oyj. pp. 891–898. Duncanson, P.L. and Toth, J.D. 2004. The truths and myths of freeze lining technology for submerged arc furnaces. Proceedings of Infacon X: Transformation through Technology, Cape Town, South Africa, 1-4 February 2004. pp. 488–499. https://www.pyrometallurgy.co.za/InfaconX/062.pdf Erwee, M.W. Reynolds, Q.G., and Zietsman, J.H. 2016. Comparison of 2D and 3D computational multiphase fluid flow models of oxygen lancing of pyrometallurgical furnace tap-holes. JOM, vol. 68, no. 6. https://www.mintek. co.za/Pyromet/Files/2016Erwee-JOM.pdf Gasik, M.I. 2013. Technology of chromium and its ferroalloys. Handbook of Ferroalloys – Theory and Technology. Butterworth-Heinemann, Oxford, UK. Chapter 8. Hancock, J.D. 2006. Practical Refractories. Cannon & Hancock, Vereeniging, South Africa. Hubble, D.H. 1999. Steel plant refractories. The Making, Shaping and Treating of Steel - Ironmaking Volume. 11th edn. The AISE Steel Foundation, Pittsburgh, PA: pp. 161–228. Inada, T., Kasai, A., Nakano, K., Komatsu, S., and Ogawa, A. 2009. Dissection investigation of blast furnace hearth—Kokura no. 2 blast furnace (2nd campaign). ISIJ International, vol. 49, no. 4. pp. 470–478. McDougall, I. 2013. Ferroalloys processing equipment. Handbook of Ferroalloys – Theory and Technology. Butterworth-Heinemann, Oxford, UK. Chapter 4. Nelson, L.R. and Hundermark, R.J. 2016. The tap-hole — key to furnace performance. Journal of the Southern African Institute of Mining and Metallurgy, vol. 116, no. 5. pp. 465–490. Pariser, H.H., Backeberg, N.R., Masson, O.C.M., and Bedder, J.C.M. 2018. Changing nickel and chromium stainless steel markets - Market review by Heinz Pariser. Proceedings of Infacon XV: International Ferro-Alloys Congress, Cape Town, South Africa, 25-28 February 2018. https://www.pyrometallurgy.co.za/ InfaconXV/0001-Pariser.pdf Steenkamp, J.D., Denton, G.M., and Hayman, D.A. 2017. Insulating or conductive lining designs for electric furnace smelting? Applications of Process Engineering Principles in Materials Processing, Energy and Environmental Technologies: A Symposium in Honor of Professor Ramana G. Reddy, San Diego, California, USA, 26 February to 2 March 2017. Springer. pp. 209–220. Steenkamp, J.D., Hockaday, C.J., Gous, J.P., and Nzima T.W. 2017. Dissipation of electrical energy in submerged arc furnaces producing silicomanganese and high-carbon ferromanganese. JOM, vol. 69. pp. 1712–1716. Steenkamp, J.D., Pistorius, P.C., and Muller, J. 2016. Insights into the potential for reduced refractory wear in silicomanganese smelters. Journal of the Southern African Institute of Mining and Metallurgy, vol. 116, no. 1. pp. 101–108. Steenkamp, J.D., Sutherland, J.J., Hayman, D.A., and Muller, J. 2016. Tap-hole life cycle design criteria: A case study based on silicomanganese production. JOM, vol. 68, no. 6. pp. 1547–1555. u
➤ In the furnaces excavated, the tap-hole was one of two high refractory wear areas, with the others being the sidewalls around the clovers of electrodes 2 and 3. The life of the tapblock in the open SAF was on average double that in the closed SAF. The results presented here do not explain the difference, and a similar study on the potential effect of refractory corrosion by gas phases present in fully enclosed SAFs will be useful. ➤ In both furnaces excavated, very little (if any) of the tapblocks remained and the SAFs were basically operated only with the quick-replaceable carbon block as the tapblock. ➤ Thermodynamic calculations indicate that both alloy and slag had the potential to corrode the tapblocks, either through dissolution or through chemical reaction. For the alloy, the main mechanism is related to dissolution of C or SiC into the alloy. For slag, chemical reactions between CrO, FeO, and SiO2 in the slag and refractory are the main mechanisms with alloy, SiC, and gas the reaction products formed. At temperatures below 1650°C, C-based refractory will potentially be the best choice in tapblock refractory, while above this temperature a SiC-based refractory would be preferred. ➤ Laboratory-scale investigations to confirm these observations will be useful from an applied research perspective, and to study the reaction mechanisms involved from a fundamental research perspective.
Figure 13 – Refractory consumption by slag for (a) C-based refractory and (b) SiC-based refractory
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Acknowledgements The support of management and personnel at a South African producer of ferrochrome and colleagues from Mintek who assisted with observations, measurements, and sampling on shift is gratefully acknowledged. The author further would like to acknowledge valuable inputs and discussions with Markus Erwee. This work is published with the permission of Mintek.
References
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Tap-hole maintenance and management on furnace 10 at Eramet Norway Porsgrunn
u
J.E. Davidsen1 and M. Honstad1 Affiliation: 1 Eramet Norway Porsgrunn, Norway. Correspondence to: J.E. Davidsen
Email:
jens.davidsen@eramet.com
Dates:
Received: 18 Mar. 2019 Revised: 20 May 2019 Accepted: 20 May 2019 Published: June 2019
Synopsis A well-maintained tap-hole channel is important in order to achieve good and stable furnace operation over time. A worn or short tap-hole channel can lead to more violent flow, significantly increasing the fume emissions during tapping and leading to more metal being mixed with and lost to the slag. Over time, a worn tapping channel might cause side channels and damage to the lining around the tap-hole area. This can result in a burn-through which is a considerable safety risk and will cause significant production losses. At Eramet Norway Porsgrunn, tap-hole maintenance has been done regularly with electrode paste for over three decades with good results. This article summarizes some of the experiences obtained from this practice on a 33 MW SiMn furnace, as well as a change to a coal tar pitch-free binder. Keywords tap-hole maintenance, lining, tap-hole wear, coal tar pitch-free.
How to cite:
Davidsen, J.E. and Honstad, M. Tap-hole maintenance and management on furnace 10 at Eramet Norway Porsgrunn. The Southern African Insitute of Mining and Metallurgy DOI ID: http://dx.doi.org/10.17159/24119717/674//2019 ORCiD ID: J.E. Davidsen https://orchid.org/0000-00020288-1556
Introduction A well-maintained tap-hole channel is important in order to achieve good and stable furnace operation over time. Metal and slag are regularly tapped from the furnace and may cause a lot of wear on the tapping channel, increasing the risk of lining damage and ultimately leading to a burn-through. On furnace 10 at Eramet Norway Porsgrunn, the tap-hole area is the part of the lining that has been by far the most prone to burn-through since the startup of the furnace. Experience shows that almost all lining problems have started with burn-through in the tap-hole area or in close proximity to the tap-hole. During tap-hole repairs, side channels with metal and slag have also been observed emanating from the main tapping channel. In addition, a shorter tapping channel will cause more violent flow. This can lead to more metal droplets being entrained in the slag, and increase the risk of the furnace not being closed properly at the end of the tap. As a burn-through is a considerable safety risk, and also causes significant production losses, systematic work has been performed over the years to maintain the tapping channel in order to prevent burn-through.
Furnace overview Furnace 10 at Eramet Norway Porsgrunn is a single-level submerged arc furnace (SAF) by Elkem Technology. The furnace was started up in 1968, but over the years several modifications and upgrades have been done to the furnace in order to increase the smelting capacity. The shell diameter and electrode diameter have, however, been kept constant. Table I show some general data for the furnace, valid after the last upgrade in December 2017. After the startup in 1968, high-carbon ferromanganese (HC FeMn) was the main product. In 1972 the furnace was switched to silicomanganese (SiMn) production and has since then been producing mainly SiMn with only shorter campaigns on HC FeMn. Eramet Norway runs the furnaces in duplex production, with slag tapped from the HC FeMn furnaces containing 30 wt.% Mn. The raw material blend used on furnace 10 has thus largely consisted of HC FeMn slag and a small proportion of ore. The ores used for SiMn production have varied through the years, usually being selected based on cost and availability. The furnace has two tap-holes situated at the front of the furnace and separated by a 30° angle. The tapping electrode, electrode B, is situated between the two tap-holes. The A and C electrode are both situated further away from the tap-holes in the ‘back’ of the furnace, giving names to the A and C The Journal of the Southern African Institute of Mining and Metallurgy
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➤
Tap-hole maintenance and management on furnace 10 at Eramet Norway Porsgrunn Table I
Table III
General data for furnace 10 after the last upgrade in December 2017
ypical alloy and slag analysis for the furnace from T July 2016 to January 2018 with standard deviation. Chemical compositions are in weight per cent. For the slag analysis: CM/S = (CaO+MgO)/SiO2 and CM/A = (CaO+MgO)/Al2O3
Transformer capacity Max. load Electrode diameter Furnace height Furnace diameter Lining type Number of tap-holes Angle between tap-holes Taps per day Average slag temperature
63 MVA 33 MW 1 700 mm 6 100 mm 11 700 mm Freeze lining with water film 2 - A and C side, single level 30° 9-12 1586 (± 15)°C
Table II
General data for the plugging paste used at furnace 10 Ingredient
Amount (wt.%)
Sand Graphite Grog Clay Phenols
50–70 5–10 10–15 10–15 < 0.1
tap-holes, respectively. Tapping is alternated between the sides, and tapping position is changed to the opposite side after each tap. The tap-hole is opened using a drill and closed with a mud gun. A standard plugging paste is used for closing the furnace, the general composition of which is summarized in Table II. The drill and mud gun are remote controlled by the tapper at a safe distance in an air-conditioned tapping booth. In the event that the tap-hole drill cannot penetrate solidified material in the tapping channel, manual oxygen lancing is used. Slag temperatures are generally in the range of 1560 to 1620°C, with an average of 1586°C for the past few years. The slag temperature is measured in the tapping stream close to the tap-hole using a dip-type thermometer. Typical composition ranges for slag and alloy are shown in Table III. Since the start up in 1968 the furnace has been out for relining for a total of seven times. Not all parts of the lining were changed in all of the outages. Table IV summarizes some of the most significant challenges and changes made to the furnace lining from startup to the present day.
Alloy
Amount (wt.%)
Slag
Amount (wt.%)
Si Mn Fe C Mn/Fe
18.4 (± 1.0) 69.3 (± 2.0) 9.9 (± 2.1) 1.7 (± 0.4) 7.4 (± 1.0)
SiO2 MnO CM/S CM/A
40 (± 1.0) 8.0 (± 2.0) 0.7–0.8 1.6–2.0
The tap-hole blocks have been pre-baked since the startup of the furnace. The design and material have varied over the years, with both carbon and SiC being tested for tap-hole block material. The observations from the plant personnel have been that SiC generally has a better resistance to chemical attack than carbon, but neither are good solutions if not maintained regularly. Since the relining in 1995 there have been only three instances of burn-through in the tapping area, all in 2009 and 2010 towards the end of the lining lifetime. This happened after the furnace was shut down due to market conditions in 2009. Following the ramp-up after this outage, the C tap side was heavily damaged due to oxygen lancing and all three burnthrough incidents occurred around this tap-hole. Since the relining in 2010 there have been no instances of burn-through or lining repair on the furnace.
Mechanisms of tap-hole wear The erosion and wear of the furnace lining and tapping channel is a complex phenomenon with many factors contributing. Some of the factors thought to have an impact in the case of furnace 10 are summarized in this section. During SiMn production the stable carbon phase will depend on the amount of silicon (Si) in the produced alloy. At lower levels of Si, below about 18 wt.%, graphite (C) will be the stable phase, while above this level silicon carbide (SiC) will be the stable carbon phase when assuming equilibrium (Olsen, Tangstad, and Lindstad, 2007). A SiMn alloy with a Si content above the coexistence point for C plus SiC can potentially
Table IV
A summary of the furnace linings, load, products, and changes made to the lining from startup in 1968 to the present Lining type
Load
Lining challenges
1968–1970
Year
Insulating
23–25 MW
Burn-through in bottom lining
Changes/comments Paste lining
1970–1976
Insulating
24–26 MW
Burn -hrough in sidewall
Paste lining
1976–1988 Insulating 25–30 MW
A lot of tap-hole repairs in the period 19841988
Only new paste sidewall lining. Tap-hole maintenance with electrode paste started in 1988
1988–1994 Insulating 28–31 MW A lot of tapping problems/high wear on lining 1994–1995 Insulating 28–32 MW Burn-through in furnace bottom 1995–2010 Insulating 25–31 MW
Three burn-throughs on the C tap-hole between 2009–2010
2010–present Freeze lining 29–33 MW
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Only changed the lining in the front and tapping area. Water-cooled tapping frame New paste lining. Installed water-cooled Cu blocks around tap-holes. Water-cooled Cu blocks. Steel pipes with cooling water welded to the furnace shell for cooling of the sidewall. New carbon block lining. Freeze lining maintained by a water film on the steel shell surrounding the lining. Water-cooled Cu blocks.
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Tap-hole maintenance and management on furnace 10 at Eramet Norway Porsgrunn dissolve carbon and precipitate SiC if it is in contact with carbon. This is caused by SiC precipitation removing carbon from the solution and moving the metal composition away from graphite saturation (Steenkamp, Pistorius, and Muller, 2016). There is little change in the point of coexistence for C and SiC with changes in temperature in the range 1550–1600°C. Cooling the area around the tapping channel will therefore have little effect on the chemical wear caused by the mechanism described above (Steenkamp, Pistorius, and Muller, 2016). Weekly averages of wt.% Si and wt.% C in the tap alloy from furnace 10 between July 2016 and February 2018 are shown in Figure 1. The data is plotted together with the calculated carbon solubility for a Mn-Fe-Si-C melt with a Mn/Fe ratio of 7.4 at 1550°C and 1600°C. The carbon solubility curves are calculated with Thermo-Calc v3.0 using the SINTEF database for the Mn-FeSi-C-Al-P system. As shown in Figure 1, the composition of the tapped alloy from furnace 10 is located mainly to the right of the coexistence point for C/SiC, indicating that SiC should be the stable form of carbon. The points are located around the solubility limit at 1550°C. Knowing from experience that the metal temperature is generally lower than the slag temperature, which averages around 1586°C, this seems like a reasonable fit. About half the points in Figure 1 are situated below the curve and thus indicate an alloy unsaturated in carbon. Such an alloys has the potential to dissolve carbon from the graphite tap-hole block, but would be less likely to dissolve a SiC block. As the tap-hole blocks in the present lining on furnace 10 are made out of graphite, chemical attack by the metal on the tap-hole block can be assumed a plausible wear mechanism. Wear can also occur through reaction between the carbon in the tapping channel and the tap slag according Equation [1]:
"
[1]
Figure 1—Weekly average of Si and C in the tap alloy for furnace 10 between July 2016 and February 2018. Calculated carbon solubility curves at 1550°C and 1600°C are plotted in the same graph, with the point of coexistence between C and SiC marked by larger circles. The red lines are calculated with ThermoCalc using the SINTEF database The Journal of the Southern African Institute of Mining and Metallurgy
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For this to occur, the activity of SiO2 in the tap slag, aSiO2, has to be greater than the equilibrium activity at a given temperature. This equilibrium activity of SiO2, assuming the activity of the other components in the reaction is unity, is calculated to be 0.17 at 1600°C andapproximately 0.48 at 1550°C (Steenkamp, Pistorius, and Muller, 2016). If a aSiO2 in the tap slag is above the equilibrium value, the reaction is assumed to be a plausible mechanism for wear of the tap-hole block. There is a large change in equilibrium activity the temperature is reduced by 50°C. Cooling the tap-hole area, for example by using water-cooled Cu blocks, will thus limit the extent of the reaction between the slag and the tapping channel (Steenkamp and Pistorius, 2015). Values of aSiO2 calculated at 1600°C for average monthly tap slag compositions for furnace 10 between July 2016 and February 2018 are shown in Figure 2. The calculations are performed with ThermoCalc v3.0 using the MTDATA MTOX database. A datapoint has been added to show the average composition from the data used by Steenkamp and Pistorius (2015). As can be seen, the calculated activity of SiO2 for furnace 10 is close to the equilibrium for the reaction described in Equation [1] at 1600°C. As the average slag temperature is below 1600°C, this should result a small driving force for the reaction between the slag and the carbon block, and thus less potential for this to be a major source of wear on the tap-hole channel. In the case of oxygen lancing being used to open the furnace, this can cause significant wear on the tapping channel. Introducing pure oxygen through a steel pipe will result in a very high temperature, and can cause both thermal and chemical wear on the carbon in the tapping channel (Erwee et al., 2015). At Eramet Norway Porsgrunn, oxygen lancing is used only when the drill gets stuck or the material blocking the tapping channel is too hard to drill through. This happens, on average, in 5 to 15% of the taps. From experience, the main reasons for having to lance the furnace open are when there is a bad fit between
Figure 2—Calculated activity of SiO2 in the tap slag from furnace 10 at 1600°C. The average composition for the tap slag shown in the graph is 8% MnO, 7% MgO, 24% CaO, 40% SiO2, 18% Al2O3, and 3% other oxides. The equilibrium activities taken from Steenkamp and Pistorius (2015) for Equation [1] at 1600°C and 1550°C, with unity activity of the other components, are shown by dashed lines. Average activity from the data presented in Steenkamp and Pistorius (2015) is shown by a red square for comparison VOLUME 119
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Tap-hole maintenance and management on furnace 10 at Eramet Norway Porsgrunn the mud gun nozzle and the area surrounding the tap-hole, or when the furnace is not completely drained and is closed while there still is a flow of metal. This causes metal to solidify in the tapping channel, and prevents the drill from penetrating all the way through. Experience with furnace operation at Eramet also shows that thermomechanical stresses can affect the lifetime of the lining in the tapping area. The large temperature changes in this area can cause thermal expansion and contraction in the material. Over time this can lead to gaps between the bricks around the tapping block, which can be infiltrated by slag and metal. Spalling might also occur if there is slag or metal infiltration in the refractories. These effects can in the long term lead to a burn-through. The implementation of water-cooled Cu blocks in this area has been shown to reduce this problem significantly as it reduces the variation in temperature and helps to solidify stray paths of metal and slag that infiltrate the lining.
Tap-hole repair and maintenance at furnace 10 The work towards improving the life of the tap-hole section of the lining has been ongoing for several decades at Eramet Norway Porsgrunn. Due to the normally slow nature of lining wear, the iterations and improvement process will also be slow. As shown in Table IV, there have been a lot of challenges with the lining and burn-through in the tap-hole area of furnace 10 over the years. The normal procedure for repairs has been to cool the area, removed as much of the damaged material as possible, and build up a new front, which is then filled by grouting. This can in some instances give good results, but the repaired area will never be as wear-resistant as the original material. Through the years, several measures have been taken to ensure as long as a lifetime as possible of the lining in the tapping area. Two of the most significant changes through the history of the furnace have been: ➤ Introduction of weekly repairs and maintenance of the tapping channel with hot electrode paste. The procedure was started in 1988 after a long period with several burnthrough events in the tapping area ➤ Installation of water-cooled Cu blocks around the tap-hole, which was started in 1994.
The procedure was established of using hot electrode paste for closing the furnace one tap on each side every week. Due to the electrode paste having a significantly higher softening point than the regular plugging paste, the paste has to be preheated prior to injection. The preheating is done at 60–70°C in a closed cabinet for two to three hours. As electrode paste emits fumes during heating, the cabinet is connected to the same off-gas suction system used to collect the tapping fumes. The heated paste is then loaded into the mud gun, and used to close the tap-hole ‘as normal’. The mud gun is kept in front of the tap-hole for 15 minutes after closing in order to allow the electrode paste to start baking. To allow for proper baking of the electrode paste in the tapping channel the next two taps are performed on the opposite tap-hole. During the tapping operation, the length of the tapping channel is visually estimated and recorded by the tapper for each tap. The data is then used by the tappers and metallurgists for following up and monitoring the health of the tapping channel. The normal drilling length is around 120 cm, and if this is decreasing noticeably over time extra maintenance with electrode paste is performed. Although not very precise, the drilling length is thought to be a good method for following the length, and thus the state, of the tapping channel on furnace 10. During the relining in 2010 the tapping block on the A-side was removed in a single piece. Figure 3 shows a picture of the inside of the tap-hole, with the original tapping block largely intact. One can clearly see both the tapping channel with the regular plugging paste, as well as the area where the tapping block has been worn away and replaced by electrode paste, as indicated by arrows. This is assumed to be sufficient proof of the effect of the regular maintenance with hot electrode paste. Figure 3 also show that the water-cooled Cu blocks have helped to keep the carbon blocks in place by reducing the movement due to thermal expansion and contraction, preventing infiltration of slag and alloy in the area.
As discussed in the previous section, Cu blocks can help to prevent chemical wear by the tapped slag. They can also help to cool down and solidify metal and slag penetrating out from the tapping channel, and will result in less variation in the temperature in the area. Chemical wear from the alloy and oxygen lancing are not countered to the same extent, and the weekly procedure with maintenance with hot electrode paste is therefore necessary to prevent the area of being worn down over time. The hot electrode paste will also fill minor cracks in the tapping channel, preventing the formation of damaging sideways channels.
Tap-hole maintenance with electrode paste on Furnace 10 The practice of maintaining the tap-hole channel with electrode paste at furnace 10 was started in 1988. The idea behind the procedure is to inject a carbon-based material which is stronger and more wear-resistant than the normal plugging paste. Electrode paste in the form of ‘A clump’, a now discontinued product delivered by Elkem Carbon, was originally used for the purpose.
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Figure 3—Left: The area around the tap-hole from the A side taken during the relining in 2010. The tapping channel with the plugging paste, the electrode paste used for maintenance, and the original tap-hole block are all visible and indicated with arrows. Right: Original drawings of the tap-hole, showing the dimensions of the Cu blocks and ta- hole block. All dimensions in millimetres The Journal of the Southern African Institute of Mining and Metallurgy
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Tap-hole maintenance and management on furnace 10 at Eramet Norway Porsgrunn Testing with coal tar pitch-free electrode paste Coal tar pitch (CTP), which traditionally has been used as a binder in electrode pastes and plugging pastes, is known to be carcinogenic (Elkem Carbon, 2017). The European Commission has decided that use of CTP will be prohibited from October 2020. To minimize the exposure of personnel to carcinogenic materials, and to be prepared for the EU prohibition, Eramet has started testing CTP-free alternatives in all procedures where CTP binders were used previously. To assess whether an electrode paste based on a CTP-free binder could be used for the tap-hole maintenance, a test was performed on furnace 10. Before starting the test, a safety analysis was performed to make sure that safety was kept as a priority. A ‘proof of concept’ test was then initiated to ensure that it would be possible to inject the CTP-free paste in a similar manner to that used for the traditional electrode paste, using the procedure described in the previous section. After it was confirmed that the CTP-free paste would inject and bake in the tapping channel, a longer test period was initiated in order to check the long-term effect of the new paste.
The test period lasted just over 5 months, starting on 17 March and ending on 27 August 2017, during which time at total of 1497 taps were carried out. The drilling length when opening the furnace was used as the primary parameter to follow wear on the tapping channel during the test. In addition, the tappers were to report if any extraordinary conditions were observed. The tapping rate was the same throughout the reference and test periods, with an average of 0.44 t/min for the alloy and 0.49 t/min for the slag. The average tap time was 46 minutes. No extraordinary conditions or observations were reported during the test. As the CTP-free paste is delivered as briquettes, the procedures had to be adapted to incorporate the difference in shape from the old A-clump. No changes were made to the heating time or temperature. Apart from this, the practical part of the test went smoothly. To validate the results for the new paste, a reference period of 1500 taps prior to starting with the CTP-free paste was selected. Historical data for the drilling length between 2011 and 2016 was also used for comparison. The distribution and averages for the drilling length from 2011 to 2016 and the reference and test periods are shown in Figure 4.
Figure 4—Distribution of the drilling length (in centimetres) for the period of 2011 to 2016 (top), reference period (middle), and the test period with CTP-free paste (bottom) The Journal of the Southern African Institute of Mining and Metallurgy
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Tap-hole maintenance and management on furnace 10 at Eramet Norway Porsgrunn As can be seen, the drilling length decreased for each of the periods, with averages of 125 cm, 120 cm, and 117 cm for 2011–2016, the reference period, and the test period, respectively. A t-test was run to compare the drilling length distribution during the reference and test periods. The test returned a P-value smaller than 0.0001, indicating that the shorter drilling length during the test period is in fact statistically significantly. Some wear of the lining in the tap-hole area is to be expected during full operation for over 7 years. However, the decrease from the historical data to the reference period and then to the test period indicates an accelerated decrease in length. Some of this decrease can most likely be attributed to an increase in the number of taps with oxygen lancing. Figure 5 shows the percentage of taps with lancing per month from 2011 to February 2018. The test period and reference period are shown by light grey and striped black bars, respectively. As the graph shows, there seems to be an increased incidence of O2 lancing from the end of 2015. Some of the increase in lancing is attributed to a worn-out drilling machine, the state of which has become gradually worse over the course of the last few years. The spike in O2 lancing for the last two months of 2017 is a result of two factors: the generally very poor condition of the drilling machine, as well as a 21-day outage for a change of the furnace transformers. The tapping equipment was completely refurbished during this outage. The decrease in O2 lancing for the two first months of 2018 is taken as a sign that the situation had started to stabilize with the new equipment after the startup of the furnace. The increase in O2 lancing is most likely the main cause contributing to the shortened drilling length for the test period. The fact that the average drilling length stabilized, although on a slightly lower level, is interpreted as an indication that the CTP-free paste works well for maintaining the tapping channel.
No other consequences have been detected after changing to the CTP-free paste. Since the test ended in August 2017, the CTP-free paste has been used regularly for the weekly maintenance of the tapping channel at furnace 10 without any signs of further decrease in drilling length. The conclusion is that the new CTP-free paste works as well as the traditional electrode pastes for maintaining the tapping channels on an industrial SiMn SAF. Currently, over 7 years after the furnace was relined, there is still no sign of wear or damage to the lining in the tapping area on furnace 10.
Conclusions The work to improve the lifetime of the lining in the tap-hole area has been ongoing at Eramet Porsgrunn since the startup of the furnace. The major contributors to wear around the tap-hole area of furnace 10 are thought to be chemical attack by the alloy, oxygen lancing, movement in the lining caused by changes in temperature, and to a minor extent chemical attack by the slag. Water-cooled Cu blocks around the tapping block has been implemented to inhibit lining movement and chemical wear from reaction with the tap slag. Regular maintenance with hot electrode paste is used to combat the wear caused by the alloy and oxygen lancing. The inspection of the lining during a dig-out in 2010 showed the tap-hole area on the A side of the lining to be almost completely intact, indicating the positive effect of the implemented actions. The change to an electrode paste with CTP-free binder was successfully carried out in 2017, reducing the exposure of personnel to carcinogenic materials.
Acknowledgements This paper is published with permission from Eramet Norway. The contributions of our colleagues Leif Hunsbedt for environmental input, Mathieu Duchamp for guidance and discussion, and Jacques Montagnon at Eramet Research for the thermodynamic calculations are greatly appreciated.
References Elkem Carbon AS, 2017. Electrode paste safety data sheet, 402/SDS/NOR, Rev. 14, 2017-10-19. Erwee, M.W., Reynolds, Q.G., Zietsman, J.H., Cromarty, R.D., and Lexmond, A.S. 2015. Towards computational modelling of multiphase flow in and around furnace tap-holes due to lancing with oxygen: an initial computational and cold model validation study. Proceedings of the Fourteenth International Ferroalloys Congress, Kiev, Ukraine 2015. pp. 174–183. https://www.pyrometallurgy.co.za/ InfaconXIV/174-Erwee.pdf Olsen, S.E., Tangstad, M., and Lindstad, T. 2007. Production of Manganese Ferroalloys, Tapir Academic Press, Trondheim. Steenkamp, J.D. and Pistorius, P.C. 2015. Tap-hole wear: analysis of daily average slag and metal compositions of a single SiMn smelter. Proceedings of the Fourteenth International Ferroalloys Congress, Kiev, Ukraine. https://pyrometallurgy.co.za/ InfaconXIV/505-Steenkamp.pdf
Figure 5—Percentage of taps with O2 lancing in the period between January 2011 and February 2018. The reference and test periods are marked with light and dark grey bars, respectively
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Steenkamp, J.D., Pistorius, P.C., and Muller, J. 2016. Insights into the potential for reduced refractory wear in silicomanganese smelters. Journal of the Southern African Institute of Mining and Metallurgy, vol. 116, no. 1. pp. 101–108.
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Multiphase flow modelling of lancing of furnace tap-holes: validation of multiphase flow simulated in OpenFOAM® M.W. Erwee1, Q.G. Reynolds2, J.H. Zietsman2, and P.J.A. Bezuidenhout2 Affiliation: 1 University of Pretoria, Pretoria, South Africa. 2 Mintek, Pyrometallurgy Division, Randburg, South Africa. Correspondence to: M.W. Erwee
Email:
Markus.erwee@gmail.com
Dates:
Received: 15 Mar. 2019 Revised: 4 Mar. 2019 Accepted: 4 Mar. 2019 Published: June 2019
Synopsis Pyrometallurgical furnaces are tapped through tap-holes that are opened with drills and oxygen lances. The lance is often used on its own or as the last step of opening the tap-hole. Interaction of oxygen with molten material inside the tap-hole can occur, and is a phenomenon that is not well understood. As part of a study that focuses on the effect of lancing on flow inside and around the furnace tap-hole, a cold-model validation was done. The validation study focuses on validating the MULTIPHASEINTERFOAM multiphase flow solver in OpenFOAM® for the problem involved. The sensitivity of the model to different materials properties (viscosity and density in particular) is presented with both cold models and the results from flow simulations in OpenFOAM®. Keywords tapping, lancing, OpenFOAM, pyrometallurgy, multiphase flow.
How to cite:
Erwee, M.W., Reynolds, J.H., and Bezuidenhout, P.J.A. Multiphase flow modelling of lancing of furnace tap-holes: validation of multiphase flow simulated in OpenFOAM®. The Southern African Insitute of Mining and Metallurgy DOI ID: http://dx.doi.org/10.17159/24119717/670/2019 ORCiD ID: M.W. Erwee https://orchid.org/0000-00015067-3439
Introduction Many pyrometallurgical operations are carried out in furnaces where molten material accumulates continuously. Once enough material has accumulated and is ready to be removed from the furnace, material is tapped through one or more tap-holes in the furnace. Tap-holes differ significantly in both geometry and materials of construction, depending on the process involved – examples of different types of tap-hole configurations can be found in, for example, Nelson and Hundermark (2014). Tapholes all share one common feature: they are high, wear areas of the furnace. Molten material at high temperature flows through these areas, after which the tap-hole is sealed, usually using tap-hole clay. The purpose of the tap-hole clay is to harden inside the tap-hole and seal it until the next tap. Once ready to tap, a pneumatic drill and/or oxygen lance is used to open the tap-hole. The extent of lancing a furnace tap-hole depends strongly on the type of furnace. In, for example, many blast furnace operations lancing is limited as much as possible, and done only if drilling was not effective. In other operations, such as on small-scale pilot furnaces with short tap-holes, only lancing is used to open the tap-hole. Lancing involves using a long metal pipe (usually steel), approximately 15 mm in diameter and varying in length, as an oxygen cutting torch. Oxygen flows through the lance and reacts with iron in the steel to form iron oxide, releasing heat. Despite significant advances in technology to enhance the safe of opening of tap-holes, lancing remains part of the process, and can be quite dangerous if not done properly. The danger stems from the fact that the process is largely manual, i.e. done by furnace staff. Both novice and expert furnace tappers can make mistakes when lancing, risking a highly focused amount of thermal energy being concentrated in the tap-hole. Aside from damage to the tap-hole, unwanted oxidation of the material inside of the furnace can occur. As well as the possible thermal disturbance, which may or may not be significant, flow around the entrance to the tap-hole inside the furnace may be affected and could contribute to wear. The work described in this paper was aimed at validating modelling work related to safer tapping practices, with a focus on oxygen lancing of furnace tap-holes. Since such a study would be virtually impossible from inside a furnace, multiphase modelling is used to study the problem in as much detail as possible. The emphasis is on some validation work done to ensure that the strengths and limitations of the model(s) can be identified for the specific problem involved – flow of gas into two fluids, which could serve as potential fluids to emulate slag and metal/alloy in real pyrometallurgical processes,
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Multiphase flow modelling of lancing of furnace tap-holes Table I
Dimensionless groups pertinent to injection of gas into two fluids (from Erwee, Reynolds, and Zietsman, 2016b) Dimensionless number
Description / purpose of the number Interfacial surface tension (σ) ratios between the top and bottom fluids as well as the gas (the fluids would represent slag, metal, and gas)
Geometry: the relative heights (h) of the top fluid (tf, e.g. slag) and bottom fluid (bf, e.g. metal) as well as the relative position of the lance with respect to the top fluid)
The ratio of viscosity (μ) of the top fluid and bottom fluid The ratio of density (ρ) of the top fluid and bottom fluid Morton number (Mo): Ratio of viscous forces to surface tension forces Reynolds number (Re): Ratio of inertial forces to viscous forces Eötvös number (Eo): Ratio of buoyancy to surface tension forces
and also to validate the model over a large range of parameters (density, viscosity, etc.).
Dimensional groups pertinent to the lancing problem Previous work by the authors (Erwee, Reynolds, and Zietsman, 2016b) revealed, by dimensional analysis, the most pertinent dimensionless groups that describe the lancing problem. These parameters allow for matching, at least in part, computational modelling work that can be done in a laboratory to the actual furnace problem. Ten dimensional groups that were identified are listed in Table I (see also Figure 1). The exact details of these groups are given elsewhere (Erwee, Reynolds, and Zietsman, 2016a, 2016b). However, they are listed here to allude to which parameters are important to the lancing problem. From the dimensional analysis, it was very clear that the physical properties of the materials involved play an important role in describing the problem, as does the level of the bottom and top fluids relative to the lance. This might seem obvious to any pyrometallurgist, but also prompts the reader to recall how
Figure 1—Parameters to consider for the lancing problem (geometry not accurate or to scale, just shown for simple orientation of fluids relative to lance) (Erwee, Reynolds, and Zietsman, 2016b)
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difficult it is to both measure and/or model the said physical properties. Furthermore, finding materials that could be used at room temperature that would satisfy the dimensionless groups involved in order to validate a computational model is equally difficult, since so many groups were identified. One group, the surface (and hence interfacial) tension ratios, is very important. For most metals and slags, the surface tension values are quite high, in the order of 0.4-1.0 N/m (Guthrie, 1989), whereas for fluids at room temperature such as oil and water, these values are quite low, e.g. the surface tension for water is approximately 7.2 × 10-2 N/m. This is just one example of differences that often make matching properties difficult. The Morton and Eötvös numbers are of importance, as they describe the shapes of bubbles/drops moving in a fluid. This would seem somewhat superficial to consider for opening a furnace tap-hole; however, when oxygen reacts with metal inside a furnace during lancing, the reaction rate is often, for example, dependent on the shape and size of the surface – i.e. to better understand this behaviour in future models (not this work), it would be important to incorporate this aspect of the physics for the fluids involved. The dimensionless numbers can be calculated for a particular furnace, using the dimensions of the lance, the slag and metal/ alloy levels, as well as the density, viscosity, and surface tension. Using this as a baseline, one can potentially source fluids at room temperature that could emulate slag and metal if some of the dimensionless groups match up. Perfect matching is not always possible, but if many different materials are used at room temperature, one can build a set of validated data for the model. Even if these do not match the real furnace fluids perfectly, if the model fits a validated data-set over a range of dimensionless numbers, it could provide insight into the limitations of the model results and allow for more confidence in extrapolations of the model parameters. The Journal of the Southern African Institute of Mining and Metallurgy
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Multiphase flow modelling of lancing of furnace tap-holes
Figure 2—Example footage from the Natsui experiment (Natsui et al., 2016)
The Natsui case
Computational model
Many computational fluid flow studies have been conducted and validated against, for example, water-based cold models. For brevity, these studies are not repeated in this paper. An excellent study that resembles a metal-slag-gas system quite closely is that by Natsui et al., (2016). The authors used a mixture of molten chloride salts, metallic tin, and argon gas to verify a smoothed particle hydrodynamics model of interfacial flow of two immiscible melts. Argon gas was bubbled through a small Pyrex beaker containing the molten material and filmed for validation of the model. An example of the behaviour observed is shown in Figure 2. Of note is the fact that a significant amount of effort went into setting up the experiment, with visual observations being crucial to match the computational model to the hightemperature model. The authors of this paper are currently setting up a similar experiment to make more observations to validate the model described later in the paper.
OpenFOAM® (openfoam.org, 2018) is an open-source framework, based in C++, in which fluid flow can be simulated using the volume-of-fluid (VOF) method (Hirt and Nichols, 1981). The specific method involves locating and tracking the fluid-fluid interface (‘free surface’) on a mesh, while fluid flow is described using the Navier-Stokes equation. The computational domain was set up on the same scale as the rectangular tank for cold model experiments, except for the total height, which was limited to 75 mm since the total fluid height was only 40 mm. The domain was meshed using the open-source tool gmsh (Geuzaine and Remacle, 2009) version 3.0.6 and then imported into OpenFOAM® version 6. The mesh was converted to a polyhedral mesh with approximately 952 000 equisized elements of 0.5 mm each. A standard solver exists in OpenFOAM® for multiphase flow problems and was used for this work. The specific solver is called ‘MUTIPHASEINTERFOAM’. The solver or model consists of an incompressible flow solver based on the Navier-Stokes and continuity equations. Furthermore, a phase separation model is included. A detailed description of the model and governing equations, with an example of their use, is given in Andersson (2010). The nature of the problem leans towards a special interface capturing scheme, developed by Ubbink and Issa (1999). The simulations were decomposed to run in parallel on the Lengau supercomputer at the Centre for High Performance Computing in Cape Town, South Africa. In this case, the simulations were run across 480 processors (30 nodes with 16 cores each and 16 MPI processes per core). On average, a case ran for approximately 6000 core hours to solve a model of 2.5 seconds in real time (time-step = 0.01 seconds). Results were processed using Paraview 5.4.1.
Cold model for this study While the set-up of a similar model to that of Natsui et al. is under way, the authors opted to conduct similar tests using a rectangular tank of 50 × 50 × 150 mm with a gas inlet at the bottom. The gas inlet had a diameter of 5 mm and was mounted flush against the bottom of the tank. Mercury, silicone oil, and water were used as fluids. Mercury was chosen since it has a high surface tension, which is important for this problem, as described earlier. The work with mercury, silicone oil (in this case a viscosity standard), and water is meant to complement work like that of Natsui et al., such that the computational model can be validated across a wide range of materials properties. The materials properties for the fluids involved in this study are given in Table II. For the cold model experiments, technical grade nitrogen gas was used and injected at flow rates between 0.2 and 0.8 L/min. The level of the top fluid (oil or water) was kept constant at 30 mm, and that of mercury at 10 mm. An Olympus i3 high-speed video camera was used to film the behaviour of the interface between the two fluids as gas bubbles rose through the tank.
Results and discussion Several cases were set up and run, but only two cases are presented in this paper. These cases showed the most salient
Table II
Materials properties at 20°C Material
Density (kg/m³)
Viscosity (Pa.s)
Surface tension (N/m, with air)
Mercury
0.485
13560
0.0015
Silicone oil
978
0.498
0.04
Water
997
8.90 x 10-4
0.07
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Figure 3—Mercury-water experiment (gas flow rate 0.8 L/min) (The time scale for the photographs is not shown, as these are snapshots of high-speed camera footage after reaching a steady state, frames are extracted from 2 seconds of real time. Each block in the background of the photographs is 10 × 10 mm)
differences between the cold model and the computational model. For this paper, the focus is on visual comparison of the behaviour of the interface and movement of the bubbles. This is because there are few other parameters that can be measured against output from the model. Previous work (Erwee, Reynolds, and Zietsman, 2016a) involved measuring pressure signals against the wall of the tank, but when fluids other than water
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are involved, this proves difficult. This is, however, still being explored by the authors. Figures 3 and 4 show the behaviour of the computational model and physical cold models for mercury-water and mercuryoil respectively. When played as video files (only frames are shown in this paper), the general behaviour of both cases seems, at first glance, to agree. When comparing, for example the case The Journal of the Southern African Institute of Mining and Metallurgy
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Figure 4—Mercury-silicone oil experiment (gas flow rate 0.8 L/min) (The time scale for the photographs is not shown, as these are snapshots of high-speed camera footage after reaching a steady state, frames are extracted from 2 seconds of real time. Each block in the background of the photographs are 10 × 10 mm)
with water (Figure 3) to that with oil (Figure 4), deformation of the mercury-water interface and water-air interface is significantly higher than in the mercury-oil case. This is despite very small differences in both density and surface tension of the water and oil. Notably, the high viscosity of the silicone oil seems to play an important role in this phenomenon, which is somewhat counterintuitive when compared to some of the cases run by The Journal of the Southern African Institute of Mining and Metallurgy
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Natsui et al., (2016). What was noteworthy when comparing the cases was the shape of the bubbles. In the mercury-water experiments, the bubbles wobbled significantly more and did not retain a spherical shape, as they did in the water-silicone cases. The same behaviour was observed in the computational model. A noticeable difference between the computational models and experiments VOLUME 119
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Multiphase flow modelling of lancing of furnace tap-holes was the behaviour of the mercury films on the interface between gas and fluid (water or oil). This difference is due to the fact that film drainage is not accounted for in the computational model. This, however, does not seem to influence, at least visually, the interfacial deformation, but will need to be explored further, along with interfacial energy, in future studies. In the experimental model it was observed that film drainage was significantly faster in the mercury-water case than in the mercury-oil case. For the mercury-oil case, it was found that the mercury film around the gas film drains off when the bubble has already left the interface, while in the water case it drained off before the bubble leaves the interface. This behaviour can be observed by comparing the bursting/draining of the mercury bubble in Figures 3f and 3g to that in Figures 3f through 3h. Similar results were obtained by Natsui et al., (2016). Another difference between the computational model and experiments is that of gas entrapment in the silicone oil. Small bubbles remained suspended in the silicone oil in the experiments, but not in the computational model. Similarly, many small beads of mercury remain on the surface of the mercuryoil interface in the experiments, but this is not observed in the computational model. This, again, is attributed to the fact that film drainage is not accounted for the in the computational model. Although film drainage is not accounted for in the computational model, the bulk behaviour, especially interfacial deformation behaviour, is modelled well. Film drainage models are inherently difficult to incorporate in computational models, but are being developed by others, at Mintek as well as the CSIR. Interfacial deformation, however, is important for the case of lancing. When oxygen gas moves through a metal or slag phase, there is potential for mixing between slag and metal. This is somewhat quantifiable by interfacial deformation, i.e. a certain degree of deformation could lead to more mixing between metal and slag and metal droplets can be entrained in the slag. Entrained metal in the slag is a direct loss to the economics of the process as metal is lost and not necessarily recovered from the slag in subsequent processes. Aside from metal entrainment, metal oxidation could also occur. Notably, no oxidation of the material is accounted for in the computational model, as it is beyond the scope of the project. This is an extremely important point, since most metals oxidize when contacted with oxygen. Metal oxidation is exothermic, which could lead to localized heating around the tap-hole area, which can be detrimental to the tap-hole and furnace. Oxidation of metal is also usually accompanied with some release of gas (highly dependent on the system), which could lead to uncontrolled tapping. Coupling CFD simulations to oxidation of metal is inherently complex (Zietsman, 2016) and will be studied in future outside of this project.
Conclusions To validate a multiphase fluid flow model for the case where gas is injected into a metal-slag mixture, cold model experiments with water, oil, and mercury were conducted. These experiments form part of a larger project on the validation of computational models in OpenFOAM®. Reasonable agreement was found between the computational models and the experiments conducted as far as bulk flow behaviour and interfacial deformation are concerned. However,
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film drainage behaviour in the experimental model is not captured in the computational model. This is expected, as film drainage models are quite complex, but for the lancing problem film drainage is not necessarily as important as interfacial deformation. The behaviour matches work done by previous authors in which molten salt-tin mixtures were used as fluids. It was found that viscosity plays a significant role in the way in which bubbles behave in the system. Further work is ongoing. Significant effort is being made to obtain more quantitative data from the experiments as well as the computational model. Furthermore, an experimental rig like the one used by Natsui et al., (2016) (molten slag-tin mixture) is being built to run more experiments.
Acknowledgements This paper is published by permission of Mintek. The contributions of our colleagues are gratefully acknowledged. The authors would also like to thank the staff of the Centre for High Performance Computing for their assistance and for making computational resources available for this project.
References OpenFOAM. 2018. The OpenFOAM Foundation. https://openfoam.org/ Anderson, P. 2010. CFD with OpenSouce Software: Tutorial multiphaseInterFoam, http://www.tfd.chalmers.se/~hani/kurser/OS_CFD_2010/patrikAndersson/ patrikAnderssonReport.pdf Erwee, M., Reynolds, Q., and Zietsman, J. 2016a. Comparison of 2D and 3D computational multiphase fluid flow models of oxygen lancing of pyrometallurgical furnace tap-holes. JOM, vol. 68, no. 6. pp. 1556–1562. Erwee, M., Reynolds, Q., and Zietsman, J. 2016b. Computational multiphase flow modelling and dimensional analysis study of oxygen lancing of pyrometallurgical furnaces. Proceedings of the 10th South African Conference on Computational and Applied Mechanics (SACAM2016). South African Association of Theoretical and Applied Mechanics (SAAM), Potchefstroom. pp. 146–154. Geuzaine, C. and Remacle, J. 2009. Gmsh: A three-dimensional finite element mesh generator with built-in pre- and post-processing facilities. International Journal for Numerical Methods in Engineering, vol. 79, no. 11. pp. 1309–1331. Guthrie, R. 1989. Engineering in Process Metallurgy. Oxford University Press, New York. Hirt, C. and Nichols, B. 1981. Volume of fluid (VOF) method for the dynamics of free boundaries. Journal of Computational Physics, vol. 39, no. 1. pp. 201–225. Natsui, S., Nashimoto, R., Kikuchi, T., and Suzuki, R. 2016. SPH analysis of interfacial flow of the two immiscible melts. Proceedings of the 10th International Conference on Molten Slags, Fluxes and Salts (MOLTEN16), Seattle, Washington, USA. Reddy, R.G., Chaubal, P., Pistorius, P.C., and Pal, U. (eds). Wiley, Hoboken, NJ. pp. 589–596. Nelson, L. and Hundermark, R. 2014. ‘The tap-hole’ – key to furnace performance. Proceedings of Furnace Tapping 2014, Muldersdrift, South Africa, 26-28 May 2014. Southern African Institute of Mining and Metallurgy, Johannesburg. pp. 1–32. Ubbink, O. and Issa, R. 1999. Method for capturing shape fluid interfaces on arbitrary meshes. Journal of Computational Physics, vol. 153. pp. 26–50. Zietsman, J.H. 2016. Efficient storage and recall of slag thermochemical properties for use in multiphysics models. Proceedings of the 10th International Conference on Molten Slags, Fluxes and Salts (MOLTEN16), Seattle, Washington, USA. Reddy, R.G., Chaubal, P., Pistorius, P.C., and Pal, U. (eds). Wiley, Hoboken, NJ. pp. 635–644. u The Journal of the Southern African Institute of Mining and Metallurgy
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Slag carry-over and the production of clean steel
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P.C. Pistorius Affiliation: Department of Materials Science & Engineering, Carnegie Mellon University, Pittsburgh, PA, U.S.A. Correspondence to: P.C. Pistorius
Email:
pistorius@cmu.edu
Dates:
Received: 1 Nov. 2018 Revised: 26 Apr. 2019 Accepted: 26 Apr. 2019 Published: June 2019
Synopsis For effective steel refining in the ladle (secondary steelmaking) the amount of steelmaking slag that is transferred from primary steelmaking must be limited, because the steelmaking slag contains iron oxide, manganese oxide, and phosphorus oxide. Adverse effects of slag carry-over include increased consumption of deoxidizers, phosphorus pick-up by the steel, and increased ladle refractory wear. Infrared imaging at longer wavelengths is an effective non-contact method to detect slag in the tapping stream. Various devices are used to shut off the tapping stream rapidly. These include slag darts, pneumatic slag stoppers, and slide gates for steelmaking converters, and eccentric bottom tapping for electric arc furnaces. Keywords steelmaking, phosphorus, slag carry-over.
How to cite:
Pistorius, P.C. Slag carry-over and the production of clean steel. The Southern African Insitute of Mining and Metallurgy DOI ID: http://dx.doi.org/10.17159/24119717/KN01/2019 ORCiD ID: P.C. Pistorius https://orcid.org/0000-00022966-1879
Introduction Oxygen converters and electric arc furnaces are used for primary steelmaking. For both process types, the product (liquid crude steel) has a similar composition: largely liquid iron, with around 800–1000 parts per million (ppm, by mass) dissolved oxygen and low but significant concentrations of carbon, nitrogen, and phosphorus. The steel is tapped into a ladle, which serves as the refining vessel to adjust the final steel composition and temperature before casting. Some steelmaking slag is inevitably tapped with the liquid steel. The mass of steelmaking slag tapped with the steel is termed ‘slag carry-over’. Controlling and limiting slag carry-over from steelmaking is a primary requirement for clean steel production (Fandrich, Lüngen, and Wuppermann, 2008). Steel cleanliness refers to limiting the concentrations of dissolved impurities (such as nitrogen, oxygen, phosphorus, and sulphur) and of nonmetallic inclusions (oxides, sulphides, and nitrides) in liquid steel and in the final solid steel product. Of these, the strongest effect of slag carry-over is on the content of dissolved phosphorus in the liquid steel. Slag carry-over increases the amount of deoxidant needed, increases wear of ladle refractory, and causes reversion of phosphorus (transfer of phosphorus from slag back to the steel). This paper briefly reviews the fundamental origin of these effects, and summarizes the practical methods that are used to limit slag carry-over.
Steelmaking process conditions In primary steelmaking, conditions are relatively oxidizing because of oxygen injection. In steelmaking converters, oxygen is injected to react with the carbon in the hot metal (containing some 4% carbon) to convert this into crude steel with around 0.1% carbon. The total oxygen consumption in steelmaking converters is 50–70 Nm3 per ton steel (Remus et al., 2013). In electric arc furnaces (EAFs), oxygen is used to lance molten scrap, and oxygen also reacts with carbon (from injected carbon and carbonbearing raw materials such as pig iron). Combustion of carbon provides energy and produces carbon monoxide to foam the furnace slag. The median oxygen consumption in electric arc furnaces in the USA is approximately 35 Nm3/t (Association for Iron & Steel Technology, 2018b). For both converters and EAFs the resulting oxygen activity in steel, expressed as the partial pressure of O2, is around 10-9 atm. The relatively oxidizing conditions result in a high concentration of FeO in the steelmaking slag (typically around 25% FeO, with some MnO present too). Oxidizing conditions and a high slag basicity are necessary to dephosphorize the steel (Turkdogan, 1996). Phosphorus, which originates from the The Journal of the Southern African Institute of Mining and Metallurgy
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Slag carry-over and the production of clean steel steelmaking raw materials, must be removed to achieve the required mechanical properties of steel. A typical specification is that the phosphorus concentration in steel should not exceed 150 ppm by mass. Phosphorus is removed from steel by oxidation, converting neutral phosphorus atoms to phosphate ions that dissolve in the slag. Primary steelmaking is the only step in steel production that allows for dephosphorization, and phosphorus removal is an important requirement in optimization of primary steelmaking practices (Wünnenberg and Cappel, 2008). The extent of dephosphorization can be expressed as the ratio of the mass percentage of phosphorus in the slag to that in the steel: [1] This ratio LP, the phosphorus partitioning coefficient, depends on the temperature, slag composition, and reaction kinetics; typical values are in the range of 50–100 for primary steelmaking. For LP = 50, every 20 kg of slag contains as much phosphorus as 1 t of steel. The much higher concentration of phosphorus in steelmaking slag than in steel is one of the main reasons why it is important to limit slag carry-over. After tapping into a ladle, the steel is deoxidized by adding elements with a strong affinity for oxygen (such as aluminium, silicon, or silicon and manganese). Oxygen must be removed to avoid the loss of (subsequently added) alloying elements by oxidation, and to avoid carbon monoxide porosity during solidification. Deoxidation lowers the concentration of dissolved oxygen from around 800 ppm to just a few ppm (in the case of aluminium-killed steel). The much lower dissolved oxygen concentration is reflected in an O2 activity of around 10-15 atm., some six orders of magnitude smaller than in primary steelmaking. The much more reducing conditions in the ladle (secondary steelmaking conditions) cause a large shift in redox reactions. At equilibrium, the concentrations of iron oxide and phosphate in the ladle slag are near zero: all the FeO, MnO, and P2O5 in carry-over furnace slag would be reduced by deoxidation and returned to the steel. The typical concentration of silica (SiO2) in steelmaking slag is 15–20% (Remus et al., 2013). Silica from carry-over slag is partially reduced by aluminium in deoxidized steel, causing some silicon pick-up by the steel.
and the sum (%FeO + %MnO) in ladle slag at the end of ladle treatment (Schwerdtfeger, 1983). Low inclusion concentrations were associated with (%FeO + %MnO) less than approximately 4%. However, there are subtleties to the effect of (FeO + MnO) on oxide inclusion composition: the less-reducing condition imparted by retaining a few per cent of FeO and MnO in the ladle slag serves to retard transfer of dissolved magnesium to the steel, and so delays formation of magnesium spinel inclusions (approximate composition MgAl2O4) (Ahlborg, 2001; Kumar, Ahlborg, and Pistorius, 2017). Spinels are associated with clogging of continuous caster nozzles, so delaying formation of spinels during ladle treatment (by retaining some FeO and MnO in the ladle slag) can be beneficial for some steel grades. Deoxidizer consumption and phosphorus reversion are strongly affected by slag carry-over, and this can be readily seen from a simple mass balance. The calculation was based on steel that contains 800 ppm dissolved oxygen upon tapping, with the steelmaking slag containing 0.25% P and a total of 25% FeO plus MnO. The steel was assumed to be deoxidized with aluminium after tapping into the ladle, with 0.05% Al remaining in solution in the steel after deoxidizing. The amount of deoxidizer (Al) required is the sum of the Al that goes into solution (0.05%, or 0.5 kg Al per ton of steel) and the Al required to reduce dissolved oxygen (from the steel) and FeO + MnO (from the slag) according to the following reactions:
Effects of slag carry-over A survey of steel plants that was published more than two decades ago (Block and Piotrowiak, 1996) remains a useful summary of the main reasons to limit slag carry-over. For both electric arc furnace and oxygen converter plants, the main reported advantages of lower slag carry-over are improved oxide cleanliness of steel, less consumption of deoxidizer, decreased reversion of phosphorus (or of silicon), and improved ladle refractory life. ‘Oxide cleanliness’ refers to the concentration of inclusions in the liquid steel. The basis of the effect of slag carry-over on oxide cleanliness is that FeO and MnO in the slag are sources of oxygen, and continue to react with dissolved aluminium (or other deoxidizers) in the steel to form alumina (or other oxide) inclusions until the concentrations of FeO and MnO are near zero. Such sustained formation of fresh oxide inclusions tends to increase the total inclusion concentration in the steel: a correlation was reported between the inclusion concentration
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Figure 1—Calculated effect of carry-over of steelmaking slag on consumption of aluminium (deoxidizer; lower graph), and on phosphorus reversion to the steel (upper graph; horizontal grey bars give the phosphorus reversion values reported by Di Napoli Guzela et al., 2003). The estimated thickness of the layer of carry-over slag is also shown in the upper graph The Journal of the Southern African Institute of Mining and Metallurgy
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Slag carry-over and the production of clean steel 2Al + 3O → Al2O3 [2a] 2Al + 3FeO → Al2O3 + 3Fe
[2b]
2Al + 3MnO → Al2O3 + 3Mn
[2c]
Reduction of SiO2 and P2O5 is not considered in the aluminium reactions – while phosphorus reversion has a significant effect on steel composition, the effect of reduction of these two oxides on Al consumption is small compared with reactions [2a] to [2c]. The calculated relationship between slag carry-over and deoxidizer consumption is shown in Figure 1, illustrating that the effect is significant. The calculated extent of phosphorus reversion is also notable; phosphorus reversion results from all the phosphorus in the carry-over slag (0.25% of the slag mass) returning to the steel upon reduction (deoxidation). As Figure 1 illustrates, the calculated values for phosphorus reversion are consistent with the industrial observations of Di Napoli Guzela et al., (2003). They reported that without a slag stopper (described later), the amount of carry-over slag from oxygen converters is 10––15 kg per ton of steel, resulting in approximately 30 ppm of phosphorus reversion; with a stopper, slag carry-over is reduced to 3–5 kg/t and phosphorus reversion to some 10 ppm. Figure 1 also shows the depth of the layer of carry-over slag on the steel in the ladle. This depth was calculated for a slag density of 2.9 t/m3 (Mills and Keene, 1987) and assuming that the ladle is approximately cylindrical, with the depth of the steel equal to the ladle diameter. As Figure 1 indicates, the depth of the layer of carry-over slag is significant (typically several centimetres); measurement of the thickness of this layer is one of the approaches used in steel plants to quantify slag carry-over. As described by Abraham and Chen (2011), the procedure is to dip a steel rod into the steel bath in the ladle after tapping: the rod melts off at the steel-slag interface and slag adheres to the length of the rod within the slag; this length is readily measured. The adverse effect of slag carry-over on the wear of ladle refractory is generally ascribed to the presence of MnO in the steelmaking slag. Experimental studies have shown that increased manganese oxide concentrations increase the wear rate of MgO-carbon refractory, which is typically used for the ladle slag line (Ikesue et al., 1988; Um et al., 2012). The corrosive effect of MnO appears to be linked to both oxidation of carbon by MnO (causing loss of carbon from the refractory) and the rapid mass transfer of MnO in the slag (Lee et al., 2001).
Limiting slag carry-over Slag carry-over is limited by detecting slag in the tap stream, and then interrupting the flow of steel and slag. In EAFs, an additional approach is to retain a significant ‘hot heel’ of steel in the furnace to serve as a barrier to slag flow. Each of these approaches is briefly described here. An alternative approach is to remove slag from the steel in the ladle after tapping, by raking off the slag. This is a less common approach (Block and Piotrowiak, 1996), and complete slag removal is difficult to achieve (Steneholm et al., 2018).
Slag detection Surveys of steel plants 20–25 years ago indicated that most plants relied on visual detection of slag in the tapping stream (da Silva, Bergman, and Lindfors, 1993; Block and Piotrowiak, The Journal of the Southern African Institute of Mining and Metallurgy
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Figure 2—Difference in spectral radiance temperature between slag and steel, for different wavelengths. Calculated for steel and slag at 1650°C. The difference is much larger in the mid-infrared (>8 µm) than in the visible range (around 0.5 µm). Adapted from Peacock (2000)
1996). At that time, the only alternative was electromagnetic detection of the slag – requiring a detector at the tap-hole, with associated maintenance concerns (Boyle et al., 2001). The electromagnetic slag detection system uses two coils around the tap stream. Alternating current applied to one coil induces an electromagnetic field that is detected by the other coil; the difference in electromagnetic properties of steel and slag (within the coils) affects the field and the difference is used to detect the presence of slag in the tap stream (Sahai and Emi, 2008; AMEPA, 2018). In contrast to the small number of plants that used a method other than visual detection of slag 20–25 years ago, some 84% of the plants listed in the latest AIST Basic Oxygen Furnace Roundup (Association for Iron and Steel Technology, 2018a) use a slag detection method (the plants surveyed are worldwide, but do not include China, Japan, Russia, South Korea, or continental Europe). It appears that the development that made reliable nonvisual slag detection possible is infrared detection of slag in the tap stream. The infrared measurement relies on the difference in emissivity between steel and slag; this difference is much larger in the mid-infrared range (several µm wavelength) than in the visible range (Peacock, 2000). One way to express this difference is with the spectral radiance temperature, which is the temperature of a black body that would give the same radiance as slag or steel, at a particular wavelength. Slag has higher emissivity than steel, and so appears brighter. Figure 2 demonstrates that the resulting difference in spectral radiance temperature is much larger at longer wavelengths; the difference remains large even with significant attenuation (absorption of light by fumes) (Peacock, 2000). This means that the infrared camera can be placed at a safe distance from the tapping stream, and can provide a reliable non-contact indication as soon as slag is detected (Viale et al., 2007). Such detection systems are available from several equipment suppliers (Luk’yanov et al., 2017).
Slag stopping To limit slag carry-over, it is necessary to stop the flow soon after slag is detected: A typical tapping rate of a large steelmaking VOLUME 119
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Figure 3—Illustration of methods to stop slag flow from an oxygen converter: (a) slag dart; (b) pneumatic slag stopper; (c) slide gate. Adapted from Orehoski and Gray (1986), Eastwood (1995), Wiesel (2003), and Maanshan Yushan Metallurgy New Materials
converter is around 1.2 t/s (Luk’yanov et al., 2017), emphasizing the need for rapid interruption of the flow as soon as slag is detected. During tapping of steelmaking converters and EAFs, the vessel is tilted to pour out the steel. One approach is to tilt back the furnace when slag is detected, but this is too slow to limit slag carry-over to low levels; for steelmaking converters, typical carry-over in such cases is 10–15 kg/t (Di Napoli Guzela et al., 2003).
Slag stopping: steelmaking converters Figure 3 illustrates methods commonly used to stop slag flow from steelmaking converters. The dart (Figure 3a) has an overall density that is between that of the slag and the steel. When it is dropped into the tap-hole, it limits slag flow through the taphole and counteracts vortex formation; if allowed to form, the vortex would draw slag through the tap-hole before all the steel has been drained. Figure 3b shows a pneumatic slag stopper; this swings into place (by the action of a pneumatic cylinder) and stops the flow by blowing gas into the tap-hole at a high supply pressure, of around 10 bar (Boyle et al., 2001). A slide gate (Figure 3c) uses sliding refractory components, each with a hole for tapping: when the holes in the components are aligned, steel can flow; when slag is detected, the components are moved to take the holes out of alignment and so stop the flow (Lewis 1885). Wiesel (2003) described the development of slide gates for converters, including approaches to limit the effects of mechanical and thermal shock on the refractory material. An advantage of slide gates is the high rate of closure. Slide gates are now available for a wide range of converter sizes (Maanshan Yushan Metallurgy New Materials, 2018).
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Figure 4—Schematic cross-section of an electric arc furnace, showing an eccentric bottom tap-hole
Figure 5—Relationship between the amount of steel retained as a hot heel and slag carry-over from an electric arc furnace. Adapted from Abraham and Chen (2011) The Journal of the Southern African Institute of Mining and Metallurgy
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Slag carry-over and the production of clean steel Slag stopping: electric arc furnaces
Eastwood, O. 1995. Metallurgical dart and method of assembly. US patent 5,451,036.
Most modern EAFs use eccentric bottom tapping (EBT): a tap-hole at the bottom of the furnace (offset from the centre of the furnace) is used to drain out the steel (Jones, Bowman, and Lefrank, 1998); see Figure 4. The tap-hole is closed with a ’flapper’ or a slide gate. The reported lifetime of an eccentric bottom tap-hole is 150–180 heats (Heinen, 1997). Some furnaces have spouts to tap steel from the side of the furnace. A siphon spout uses a submerged entry to allow separation of steel and slag; the reported life is some 300 heats (Heinen, 1997). A slide gate can also be added to the spout to allow rapid shut-off of tapping (Jones, Bowman, and Lefrank, 1998; Heinen, 1997). In the case of eccentric bottom tapping, retaining some steel in the furnace is essential to avoid slag carry-over by vortex formation. The retained steel is termed a ‘hot heel’. The volume of the hot heel has a strong effect on slag carry-over; Figure 5 illustrates that a hot heel of some 10% of the tap mass was sufficient to decrease slag carry-over to around 7 kg/t. In Figure 5, a negative hot heel implies that there was no steel left in the furnace and the tap steel stream consisted entirely of slag at the end of tap (S. Abraham, personal communication).
Fandrich, R., Lüngen, H.B., and Wuppermann, C.-D. 2008. Actual review of secondary metallurgy. Revue de Métallurgy – CIT, vol. 105. pp. 364–374.
Conclusions Slag-free tapping is the starting point for production of clean steel. Slag carry-over of a few kilogram of slag per ton of steel is achievable with the current technological solutions: a combination of slag detection, and a method to stop flow rapidly.
Acknowledgements The support of the industrial members of the Center for Iron and Steelmaking Research at Carnegie Mellon University is gratefully acknowledged. I am also grateful for helpful comments by Mr Herman Moggee of RHI-Magnesita, and Dr Sunday Abraham of SSAB Americas.
References
Heinen, K.-H. 1997. Elektrostahlerzeugung, 4th edn. Verlag Stahleisen. Düsseldorf. Ikesue, A., Yamamoto, H., Shikano, H., and Hiragushi, K. 1988. Corrosion mechanism of MgO-C refractories by manganese oxide containing slag. Ceramic Transactions, vol. 4. Advances in Refractories Technology. Fisher, R.E. (ed). The American Ceramic Society, Inc., Westerville, OH. pp. 464–488. Jones, J.A.T., Bowman, B., and Lefrank, P.A. 1998. Electric furnace steelmaking. The Making, Shaping and Treating of Steel. 11th edn. Steelmaking and Refining Volume Fruehan, R.J. (ed.). AISE Steel Foundation, Pittsburgh, PA. pp. 525–660. Kumar, D., Ahlborg, K.C., and Pistorius, P.C. 2017. Application of kinetic model for industrial scale ladle refining process. Proceedings of AISTech2017. pp. 2693–2706. Lee, M., Sun, S., Wright, S., and Jahanshahi, S. 2001. Effects of transition metals on the kinetics of slag-refractory reactions. Metallurgical and Materials Transactions B, vol. 32. pp. 25–29. Lewis, D.D. 1885. Stopper for ladles. US patent 311,902. Luk’yanov, S.I., Suspitsyn, E.S., Pishnograev, R.S., and Krasilnikov, S.S. 2017. Survey of melt stream infrared radiation parameters at various stages of steel tapping from basic oxygen furnace. International Journal of Advanced Manufacturing Technology, vol. 88. pp. 595–602. Maanshan Yushan Metallurgy New Materials. 2018. http://en.masysyjxcl.com Mills, K.C. and Keene, B.J. 1987. Physical properties of BOS slags. International Materials Reviews, vol. 32. pp. 1-120. Orehoski, M.A. and Gray, R.D. 1986. Ladle refining processes. Iron and Steel Engineer, vol 63, no. 1. pp. 40–52. Peacock, G.R. 2000. Thermal imaging of liquid steel and slag in a pouring stream. Thermosense XXII. Dinwiddie , R.B. and LeMieux, D.H. (eds). Proceedings of SPIE, vol. 4020. pp. 50–60. Remus, R., Monsonet, M.A.A., Roudier, S., and Sancho, L.D. 2013. Best available techniques (BAT) reference document for iron and steel production. European Union, Luxembourg. http://eippcb.jrc.ec.europa.eu/reference/
Abraham, S. and Chen, S. 2011. Process for optimizing steel fabrication. US patent application 2011/0174457 A1.
Sahai, Y., and Emi, T. 2008. Tundish Technology for Clean Steel Production. World Scientific Publishing, Hackensack, NJ.
Ahlborg, K. 2001. Seven ways to shut down the caster. Proceedings of the 84th Steelmaking Conference, Baltimore, MD, 25-28 March 2001. Iron & Steel Society of AIME. pp. 861–869.
Schwerdtfeger, K. 1983. Present state of oxygen control in aluminum deoxidized steel. Archiv für das Eisenhüttenwesen, vol. 54. pp. 87–98.
AMEPA Process Measurement Systems. 2018. ESD – Electromagnetic slag detection. https://amepa.de/en/produkte/tsd-slag-detection/
Steneholm, K., Andersson, N.A.I., Tilliander, A., and Jönsson, P.G. 2018. The role of process control on the steel cleanliness. Ironmaking & Steelmaking, vol. 45. pp. 114–124.
Association for Iron & Steel Technology. 2018a. AIST basic oxygen furnace roundup. Iron & Steel Technology, vol. 15, no. 4. pp. 194–201. Association for Iron & Steel Technology. 2018b. AIST electric arc furnace roundup. Iron & Steel Technology, vol. 15, no. 1. pp. 124–147. Block, F.-R., and Piotrowiak, R. 1996. Verringerung der Menge an Mitlaufschacke bei der Stahlherstellung. Stahl und Eisen, vol.116. pp. 95–99. Boyle, C.A., Olszewski, J.A., DiFiore, N., and Daum, T. 2001. Installation of pneumatic slag stoppers at U.S. Steel’s Edgar Thomson Plant. AISE Steel Technology, May. pp. 39–40.
Turkdogan, E.T. 1996. Fundamentals of Steelmaking. The Institute of Materials, London. Um, H., Lee, K., Choi, J., and Chung, Y. 2012. Corrosion behavior of MgO-C refractory in ferromanganese slags. ISIJ International, vol. 52. pp. 62–67. Viale, M., Martin, O., Muratori, F., Bertezzolo, U., Perez, J., and Usart, J. 2007. Application of on-line infrared thermography in steel making industry. Thermosense XXIX. Knettel, K.M., Vavilov, V.P., and Miles, J.J. (eds_. Proceedings of SPIE, vol. 6541. pp. 65410H-1–11.
Da Silva, W.C., Bergman, D., and Lindfors, N.O. 1993. Slag carryover in oxygen converters: an international review. Steel Times International. pp. 91–95.
Wiesel, M. 2003. Refractories for flow control – new systems, new solutions. Refractories Engineering, March. pp. 2–8.
Di Napoli Guzela, D., de Oliveira, J.G., Staudinger, G., and Müller, J. 2003. The ultimate LD steelmaking converter. Steel Times International, April/May. pp 20–26.
Wünnenberg K., and Cappel J. 2008. Cost-saving operation and optimization on metallurgical reactions in BOF practice. Iron & Steel Technology, vol. 5, no. 11. pp. 66–73. u
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REVITALISING EXPLORATION
ACTIVITY IN SOUTHERN AFRICA POTENTIAL FOR EXPLORATION
Date: 11-12 September 2019 Venue: Misty Hills Country Hotel & Conference Centre, Muldersdrift, Johannesburg
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Mining Companies, Exploration companies, industry funders, government, academics, geologists, mining executives, consultants, students
The objective of this conference is to understand the barriers that exist in terms of establishing exploration activity, and to develop solutions in terms of opportunities, funding, skills, technology and policy.
ABOUT THE EVENT:
TOPICS:
Exploration is a critical component of the development and sustainability of the minerals industry, in any resource endowed country. Southern Africa still has vast untapped mineral resources, yet the level of exploration is at an all-time low.
• • • • • • •
Policies and Legal Issues Brownfields/Greenfields Exploration What are the potential successes Strategic/Local Success Technology Data Modelling Financing and Funding
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Managing the tap-hole life-cycle at five submerged arc furnaces producing silicomanganese at Transalloys J.J. Sutherland1 and J.P. Gous1 Affiliation: 1 Transalloys, Emalahleni, South Africa. Correspondence to: J.J. Sutherland
Email:
Kobuss@transalloys.co.za
Dates:
Received: 12 Apr. 2019 Revised: N/A Accepted: N/A Published: June 2019
Synopsis Transalloys is a silicomanganese (SiMn) producer located in South Africa, and uses only local manganese ores to produce SiMn alloy. The plant operates five open submerged arc furnaces (SAFs) with an annual capacity of 180 000 t of saleable SiMn. Maintaining tap-holes is critical for effective furnace operation, allowing proper drainage of the furnace with minimal operator interference. Using a claygun and drill arrangement installed at each furnace, tapping occurs every four hours. Tap-holes are maintained through mickey replacement and brick repairs. Three furnaces have SiC tap-holes and two have graphite tap-holes. This paper gives a review of SAF operation, furnace and tap-hole design, daily tapping operation, and maintenance practices for repairing tap-holes. Keywords tap-hole, life-cycle, submerged arc furnace, silicomanganese, Transalloys.
How to cite:
Sutherland, J.J. and Gous, J.P. Managing the tap-hole life-cycle at five submerged arc furnace producing silicomanganese at Transalloys. The Southern African Insitute of Mining and Metallurgy DOI ID: http://dx.doi.org/10.17159/24119717/KN02/2019 ORCiD ID: J.J. Sutherland https://orchid.org/0000-00020738-8090
Introduction Defining tap-hole life cycle, Steenkamp et al., (2016) highlighted four main steps: installation during furnace relines, day-to-day operations, tap-hole maintenance, and tap-hole repair. The authors argued that the design of the tap-hole area should allow for all four stages in the tap-hole life-cycle. Design principles for each stage were presented using silicomanganese (SiMn) production as a case study. Here we present a further analysis of the case study by providing practical examples of how the tap-hole life-cycle is managed at Transalloys during the design, operation, and maintenance of submerged arc furnaces (SAFs) applied in the production of SiMn.
Background Transalloys is the largest producer of SiMn in Africa. Its smelter complex is based outside the town of eMalahleni, in the Mpumalanga Province of South Africa. Transalloys was commissioned in the mid1960s as a high- and low-carbon ferrochromium plant based on the Perrin process. In 1967 the plant was converted to SiMn production because of constraints in the ferrochrome market (Basson, Curr, and Gericke, 2007; Bezemer, 1995). Today, the installed production capacity is 180 000 t of SiMn annually. Plant operations consists of five SAFs (see ratings in Table I), operating 24 hours a day, 365 days a year (including maintenance), involving 280 permanent employees, with up to 120 contract employees on site at any given time. Table I
Power ratings per furnace at Transalloys Furnace
F1
F3
F5
F6
F7
MVA rating
15
22
48
30
48
MW
14
14
28
16
28
In the SAFs (see Figure 1 for a layout), SiMn is produced by carbothermic reduction of manganesebearing ore. Manganese ore is sourced from the Kalahari Manganese Field in the Northern Cape Province. Quartz is sourced locally from South African producers. The main source of carbon is bituminous coal from South African coal mines, with imported coke also used. The specification of the SiMn alloy produced is 65% Mn, 16% Si, < 2% C, with Fe making up the balance. SiMn alloy is used as an alloying component to produce many different grades of steel. The Journal of the Southern African Institute of Mining and Metallurgy
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Managing the tap-hole life-cycle at five submerged arc furnaces producing silicomanganese at Transalloys Table III
Summary furnace refractory and tap-hole design and timelines for tap-hole repairs
Figure 1—Conceptual layout of an open submerged arc furnace producing silicomanganese at Transalloys (sketch provided by Joalet D Steenkamp)
Furnace
Furnace refractory design
Tap-hole design
Tap-hole cooling
Claygun type
1
Insulating lining
SiC bricks
Natural air cooling
Pedestalmounted
3
Semi-freeze lining
SiC bricks
Forced channel water cooling
Rail-mounted
5
Semi-freeze lining
Graphite block
Forced channel water cooling
Pedestalmounted
6
Insulating lining
SiC block
Natural air cooling
Pedestalmounted
7
Semi-freeze lining
Graphite block
Forced channel water cooling
Pedestalmounted
The SAFs are open to the atmosphere and process off-gas is combusted on top of the burden, from where it is cooled and cleaned through a baghouse system before venting to atmosphere. A mix of slag and alloy is cascade-tapped into alloy ladles and slag pots from a single tap-hole every four hours. Alloy and slag are separated post-tapping by cascade tapping and then removing slag from the alloy ladle with a slag scraper. Alloy is then layer cast in casting beds and slag is discarded onto slag dumps. The slag-to-alloy ratio is typically 1.0 to 1.3. Typical slag compositions are shown in Table II. In the submerged arc furnace operation all three electrodes are covered by raw materials. Manganese enters the process as Mn2O3 and Mn3O4 and is prereduced in the upper layers of the furnace to MnO by CO gas permeating through the furnace burden. Final conversion of MnO to Mn and SiO2 to Si is facilitated by the coke bed below each electrode, as seen in Figure 1. Table II
Typical tapped slag compositions (per cent by mass, 2669 samples with XRF) MnO
SiO2
MgO
CaO
FeO
Al2O3
B-ratio
Av.
10.73
43.28
7.31
31.94
0.41
4.31
0.91
Std dev.
±1.55
±0.92
±0.33
±1.17
±0.14
±0.30
±0.04
Why silicon carbide and carbon tap-holes?
Managing the life-cycle In this section the tap-hole designs are discussed. Table III provides a summary of the furnace refractory design and tap-hole design of each furnace. A ‘semi-freeze’ lining at Transalloys refers to two concepts used in a single furnace. Different areas in the furnace are built either as a freeze lining or an insulating lining as follows. ➤ The sidewall areas around the tap-hole are graphite bricks with graphite tiles against the shell – this is the classic freeze lining concept. ➤ The rest of the sidewalls have a graphite tile against the shell and then a 60 Alumina brick as a working lining – a standard insulating lining. ➤ On the hearth two layers of carbon blocks are installed; but beneath these block are nine layers of high-alumina bricks (three layers each of 85 Alumina, 90 Alumina, and Tab Alumina) - a combination of freeze and insulating lining.
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➤ On top of the two layers of carbon blocks a 600 mm carbon ramming is also installed – standard lining design for ferroalloy furnaces.
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Alloy produced at Transalloys is either C-saturated or SiCsaturated depending on the silicon grade of the alloy (see Figure 3) (Steenkamp et al., 2016). During excavation of a 48 MVA submerged arc furnace at Transalloys, two high-wear areas were found: the tap-hole area and the furnace hearth (Gous et al., 2014). Carbon-based refractory material formed the hot face refractory. Analysis of daily average tapped slag and metal compositions in the four months prior to excavation confirmed the potential for refractory wear through C dissolution in the metal (see Figure 3) and SiC formation by chemical reaction with SiO2 in the slag.
Design All five furnaces have one single-level tap-hole. Metal and slag are tapped simultaneously every four hours. The furnace crucible is 3–4 m deep and the tap-holes are 200 mm above the hearth floor. The Journal of the Southern African Institute of Mining and Metallurgy
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Managing the tap-hole life-cycle at five submerged arc furnaces producing silicomanganese at Transalloys In 2008 only one furnace had a proper claygun and drill, and frequent tap-hole failures were experienced on all furnaces. Tap-holes were lanced open with oxygen and closed by hand using a ‘dolly’ (see Figures 10a and b). In an effort to improve tap-hole life two design philosophies were introduced as shown in Table III, mainly to address tap-hole problems. The first was a semi-freeze lining design where tap-hole cooling was crucial, and the second an insulating design using high-wear refractory. Both these designs stemmed from a need to ensure better tap-hole life as well as easy repair once a tap-hole failed.
Semi-freeze lining design Figure 3—Carbon solubility in Mn-Fe-Si-C alloys with a Mn:Fe mass ratio of 4.47 at 1600°C, calculated with FactSage 6.4 (FSstel database); compositions given as mass percentages. For silicon contents below the inflection in the curve, the stable solid phase at saturation is graphite; SiC is the stable phase at saturation for higher silicon contents. The symbols show reported daily average silicomanganese compositions over a four-month period (Steenkamp et al., 2016)
Furnaces 5 and 7 with the semi-freeze lining design are designed with a graphite tapblock. Figure 4 shows a typical graphite tapblock design consisting of a carbon block (600 mm × 600 mm × 800 mm), two graphite lintels (350 mm × 300 mm x 800 mm), and three sacrificial mickeys (600 mm × 600 mm × 170 mm) on the working face with carbon bricks surrounding the tapblock.
Figure 4—Schematics of graphite tapblock design (redrawn from original drawings and photographs of actual installation by Driaan Bezuidenhout)
Figure 5—(a) Inside of tap-hole showing the staggered design of graphite bricks and a SiC brick tap-hole; (b) sacrificial carbon ramming on hearth and carbon taphole wall before sidewall ramming is installed; (c) former for sidewall ramming installed; (d) carbon tap-hole showing carbon tap-hole block and carbon lintel; (e) outside view of carbon tap-hole showing carbon mickey installed; (f) water cooling channels on furnace shell, front mickey, and launder flange with curved launder design The Journal of the Southern African Institute of Mining and Metallurgy
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Managing the tap-hole life-cycle at five submerged arc furnaces producing silicomanganese at Transalloys Figure 5a shows the staggered design on the sidewalls of furnace 3, allowing for optimal contact between the sidewall carbon ramming (Figure 5c) and the carbon bricks surrounding the tapblock. In the semi-freeze lining design the tapblocks, lintels, and mickeys are surrounded by carbon bricks (Figures 5b and d), allowing for maximum heat removal through external water cooling channels as shown in Figure 5f. The outside of a carbon tap-hole showing the installed mickey and the launder flange is shown in Figure 5e. Figure 5b shows the installed carbon ramming on the furnace floor (800 mm thick). The carbon ramming is taken up to 200 mm below the tap-hole. When completed, a sidewall ramming is installed (800 mm wide) as shown in Figure 5c, covering the tap- hole graphite bricks. A drill depth of 2 m is required to open a tap-hole. Figure 5f shows the tapping launder installed in position; notice a slight bend 2 m from the tap-hole. This allow for the tappers to lance and poke the tap-hole in a direct line with the
tap-hole to prevent unnecessary damage to the carbon block. Figure 5d shows the front sacrificial mickey and flange as installed. The tapping launder is bolted onto the tap-hole flange.
Insulating lining design Furnaces 1 and 6 have an insulating refractory design with no external water cooling on the furnace shell. In this design the tap-hole is constructed with SiC blocks, SiC bricks, and a SiC lintel. Figure 6 shows a typical SiC tap-hole design consisting of a two SiC blocks (600 mm × 600 mm × 600 mm) and two SiC lintels (300 mm × 800 mm × 300 mm). Lintels are installed on the inside of the tap-hole to maintain the integrity of the sidewall. When the tapblock wears only the tap-hole block can be removed with careful demolition and then a new block inserted (similar to the carbon block installation in Figure 12). Figure 7a shows the square block tap-hole design (furnace 1). The tapblock area consists of SiC bricks, whereas the rest of the furnace wall consists of SiO2/Al2O3 bricks. Figure 7b shows
Figure 6—Schematics of SiC tap-hole design (redrawn from original drawings and photographs of an actual installation by Driaan Bezuidenhout)
Figure 7—(a) Square SiC tap-hole design inside carbon bricks; (b) inside of tap-hole showing the staggered design of SiC bricks; (c) sidewall carbon ramming being installed; (d) outside view of SiC bricked tap-hole
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Managing the tap-hole life-cycle at five submerged arc furnaces producing silicomanganese at Transalloys a staggered design of the bricks on the sidewall of Furnace 6, allowing for improved contact between the sacrificial carbon ramming wall (not yet installed) and the SiC bricks around the tap-hole. Figure 7b shows the sacrificial carbon ramming on the furnace floor (600 mm thick). The ramming is taken up to 200 mm below the tap-hole. A steel pipe is inserted into the taphole in order to keep the tap-hole free from any obstructions and to facilitate the opening of the first tap after start-up. When completed, a sacrificial sidewall ramming is installed (600 mm wide) as shown in Figure 5c covering the SiC bricks forming the tapblock. A drill depth of at 1.8 m is required to open a tap-hole. Figure 7d shows configuration of bricks on the outside of the tap-hole. The face plate and tap-hole flange where the tapping launder will be bolted onto are shown.
Hybrid design Furnace 3 also has a semi-freeze lining design with external water cooling but with a SiC tap-hole. This tap-hole proved to be the most durable and only minor repairs were required in the eight years since the furnace was relined in 2008.
Tap-hole operations day-to-day All furnaces are equipped with a separate hydraulic claygun and rill supplied by Dango & Dienenthal SA (Pty) Ltd. Figure 8a shows a pedestal-type claygun and drill, four of which are installed at Transalloys, and Figure 8b a rail-type claygun and drill, only one of which is in use. Drill bits used to open tapholes are 76 mm in diameter and on average the drill depths are between 1.5 and 2 m. This section will focus on the opening of tap-holes, tapping of furnaces, closing of the tap-hole, and general tap floor operation and equipment.
Opening of tap-hole Furnaces are tapped six times a day at fixed intervals of four hours. Each tap lasts approximately 30 minutes. A hydraulic tap-hole drill is used to drill open the tap-hole. The successful opening of a tap-hole depends largely on the proper closing after the previous tap. The tap-hole is considered to be properly plugged when the entire length of the tap-hole is filled with taphole clay and no slag or metal is frozen inside the tap-hole. Before drilling the tap-hole, several important aspects need to be inspected and verified to reduce the risk of anything going wrong during the tapping process. Firstly, the drill bit is inspected to consider whether it is still sharp enough to ‘cut’ the tap-hole open as opposed to only ‘pushing’ the tap-hole open. Secondly,
the air flow through the centre of the extension rod is tested to confirm there is adequate air flow to blow out the debris that is formed during the drilling process as well as to assist in the cooling of the drill bit and extension rod during drilling. It is important that the drilling debris be blown out of the tap-hole – if this is not the case the drill bit can get stuck. Thirdly, and most critical, the alignment of the drill with the tap-hole must be checked. If this alignment is not perfect a ‘new’ tap-hole will be drilled and at the end of the tap it will not be possible to close the tap-hole as the claygun will most probable not be aligned with the ‘new’ tap-hole In case of the tap-hole not opening by drilling, an oxygen lance is used to open the tap-hole. Lancing of the tap-hole should be minimized due the fact that oxygen lancing reduces the life of the tap-hole considerably by oxidizing the refractory material. Lancing the tap-hole should be carried out on the same level and in the same direction as the drill. This is to prevent the creation of secondary tap-holes, especially towards the back of the tap-hole. The design of the tap floor should be such that a person of average height holding the oxygen lance in a comfortable position should be able to lance horizontally. The launders are therefore slightly bent approximately 2 m from the tap-hole to enable the tapper to stand directly in front the taphole when lancing. Only moderate oxygen pressure is required to lance through solidified slag and metal. If the pressure is too high the tap-hole wear will be increased dramatically due to increased oxidation. Additionally, due to the large amount of oxygen blown into the tap-hole there is a large amount of gas coming out of the tap-hole. This gas serves as a carrier gas for pieces of slag, alloy, or raw materials and can create dangerous situations with the particles moving as projectiles through the air posing the potential to cause serious injuries to persons standing on the tap floor. However, when the oxygen pressure is too low the opening of the tap-hole is very time-consuming and slag and alloys are not blown out of the tap-hole effectively. The metal and slag then solidifies in the tap-hole channel and an excessive amount of oxygen will be required to open the tap-hole channel. Oxidation then occurs towards the front of the tap-hole. This wear can lead to rat-hole formation in the sidewall if the wear is not controlled properly. In this case the tap-hole cannot be drilled open. If drilling is attempted, utmost care must be taken that the drill bit does not deviate from the normal path when drilling into the metal, for example when the drill extension moves upwards in the tap-hole during drilling. This causes ‘new’ holes to be drilled
Figure 8—(a) Pedestal claygun and drill; (b) rail-type claygun and drill The Journal of the Southern African Institute of Mining and Metallurgy
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Figure 9â&#x20AC;&#x201D;(a) Primed claygun; (b) claygun nozzle
and perforate the tapblock, reducing the life of the tapblock dramatically. Once the tap-hole is open, a good practice to maintain a good flow of alloy and slag is to use a poker bar to keep the tap-hole open and the flow unrestricted. A poker bar is also a good tool to remove pieces of raw material or electrode that get stuck in the tap-hole.
Managing alloy and slag flow If the alloy and slag flow is strong and uninterrupted, minimal intervention by the tappers is required. Occasionally the flow is slowed down by raw material mix obstructing the tap-hole. A poker bar is then used to clear the obstruction. Bigger obstructions like electrode pieces, bricks, and hardened tap-hole clay are sometimes difficult to remove with a poker bar. Then oxygen lancing is required to enlarge the tap-hole diameter to remove the obstacles. It is important to remove the obstacles continuously to ensure proper drainage of the furnace and to allow for proper closing of the tap-hole. If obstacles are not removed from the tap-hole, the plugging of the tap-hole will be inefficient. At the next tap the tap-hole will be full of frozen alloy and slag and excessive lancing will be required. Excessive lancing increases tap-hole wear and in extreme cases the integrity of the tap-hole will be compromised when sidewall bricks start to collapse into the tap-hole. The furnace is an integrated system and poor electrode management can adversely affect the life of a tap-hole, due to broken-off electrode pieces blocking the tap-hole and requiring excessive lancing. Operators need to understand that everything is integrated, therefore good electrode management is of utmost importance to prevent electrode tip losses that will cause blockages in the tap-hole.
Closing of tap-hole For effective closing of the tap-hole after each tap the full length of the tap-hole should be filled with tap-hole clay to ensure proper contact between the tap-hole sidewalls and the new tap-hole paste. This will prevent the tap-hole from opening unexpectedly between taps due to pressure from the slag and metal inside the furnace pushing the tap-hole clay out of the tap-hole or leaking between the tap-hole clay and the tap-hole sidewall. Clayguns should always be fully loaded and properly prepared before the tap-hole is opened, as shown in Figure 9a. The aim is for the clay to immediately push into the tap-hole as soon as the claygun piston starts pushing forward to displace metal and slag. The preparation of the claygun is aimed at preventing alloy
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from damaging the nozzle tip and to create a proper seal between the tap-hole and the claygun tip. The procedure for the proper preparation of the clayguns is described below. During the loading and preparation of the claygun the taphole clay in the claygun barrel must be compressed to remove any voids. This is done by pressing the claygun against the tap-hole (when closed) and pushing the cylinder forward until clay is pushed out between the claygun tip and the tap-hole. This indicates that the tap-hole clay in the claygun is fully compressed. Omitting this increases the risk of burning the claygun nozzle when closing the tap-hole against a full stream of alloy. When closing the tap-hole against a full stream of alloy there will be a period when the cylinder is pushing clay forward but no clay is coming out of the nozzle and into the tap-hole, while voids in the barrel are being filled. Even if this is only for a few seconds when the furnace is tapping strongly the nozzle will burn and then the tap-hole will not be closed effectively. When the tap-hole is not closed the furnace will continue tapping, causing possible damage to downstream equipment and injuries to people. It is good practise to always be prepared for the worst case and this simple act performed in a disciplined manner can prevent damage when it is least expected. These types of emergencies can include launder penetration, ladle penetration or tapping over the launder, electrode breaks, or many other risks in the day-to-day operation. As discussed, the alignment of the claygun and drill with the tap-hole is critical and must be checked on a daily basis. The nozzle tip needs to fit tightly around the centre of the tap-hole, ensuring that the full clay stream enters the tap-hole without squeezing out at the sides of the claygun tip. Three types of CTPV-free tap-hole clays are used at Transalloys, namely low cost (low quality), intermediate cost (standard quality), and high cost (high quality). A summary of the chemical compositions of the different types of clay is given in Table IV. Only the intermediate- and high-cost types of clay can be used to rebuild a damaged or worn tap-hole. To control operational costs low-cost clay is used in normal dayto-day operations with the expectation there will be some wear of the tap-hole. Good tap-hole practise can delay this wear. The tap-hole clay quality affects not only the tap-hole integrity, but all the aspects around the tap-hole operations. When it is noted the tap-hole is starting to wear beyond a certain limit a switch will be made to the intermediate-cost tap-hole clay with the aim of restoring the tap-hole condition. If serious tap-hole problems occur, i.e. when bricks in the tap-hole fall out, the expensive taphole clay will be used to repair the damage. A practical approach to tap-hole management must consider the cost of the tap-hole clay and the integrity of the tap-hole. The Journal of the Southern African Institute of Mining and Metallurgy
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Managing the tap-hole life-cycle at five submerged arc furnaces producing silicomanganese at Transalloys Table IV
Chemical composition of different types of tap-hole clay used at Transalloys Al2O3 (%)
SiO2 (%)
SiC (%)
Low cost
5.9
77.2
0
Moderate cost
16
69
0
63.5
18.7
14.5
Expensive
Tap floor activities Launders are cleaned after every second tap and re-dressed with river sand (see Figure 10g). Launder sand protects the carbon ramming on the launder and also makes it easy to remove the slag and alloy from the launder. Figure 10a shows some of the typical launder cleaning tools used on a tap floor. A bent oxygen pipe is used to remove alloy and slag buildups after they have been loosened with a steel roundbar. In the case of the claygun not being operational a dolly is used to close the tap-hole manually. Figure 10a also shows a former used to create a tap-hole clay plug to be used when closing the tap-hole. The steel launder is normally lined with used refractory bricks of any quality used at Transalloys (alumina, SiC, or carbon bricks, Figure 10b), then covered with carbon ramming material and finally covered with a layer of sand (see Figure 10c). The sand makes the cleaning of the launder easier in that it is easier to insert a steel ‘gwala’ below the metal/slag scull through the sand layer and lift the skull out of the launder. Launders are angled to allow for lancing and poking to be conducted parallel with the tap-hole. A launder tip is attached to the front end of the launder (see Figure 10d). High wear-resistant castable refractory is used in the launder tip to withstand the
eroding forces, especially from the alloy. The launder tip is fastened to the launder with wedges to allow for movement or a default line in case a crane driver accidentally knocks the launder tip when removing or placing the ladle. No damage to the launder or launder flange at the furnace shell will occur if this happens. Worn-out launder tips are replaced every three to four months to prevent alloy from damaging the top ring of alloy ladles.
Tap-hole and hearth maintenance programme Monitoring of furnace hearth temperatures, visual inspections of tap-holes, drill depth, tap-hole diameter, and tapping conditions are used to determine a maintenance and tap-hole repair schedule. Previous excavations and lining failures indicated that the most problematic area in the SiMn production process is the wear of the hearth refractory in the clover area (PCD) under the furnace electrodes. Subsequently, all furnaces were equipped with several hearth thermocouples to monitor hearth temperatures. Table V shows typical life-cycles of relines and tap-hole repairs for each furnace at Transalloys. Time intervals listed in Table V may be shorter in cases of catastrophic lining failures or excessive O2 lancing that will reduce the lifetime of a tap-hole. Long downtimes (in excess of 3 months) can also lead to lining failures due to uneven contraction and expansion of the hearth and sidewall refractories. Telltale signs that indicates that a tap-hole rebuild is imminent are the following: refractory bricks coming out of tap-hole, furnace tapping raw material mix at regular intervals, increase in tap-hole clay usage, and tap-hole being drilled open within less than 1 m of drill penetration. Tap-hole life can be extended by using expensive high-quality clay containing SiC (see Table III). This can buy time to properly plan for a tap-hole rebuild or repair.
Figure 10—(a) Tap-hole clay former. hook bend from redundant oxygen lance, poker bar, shovel, and dolly to close tap-hole manually; (b) redundant refractory bricks used on base of launder; (c) river sand used to dress launder; (d) removable launder tip The Journal of the Southern African Institute of Mining and Metallurgy
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Managing the tap-hole life-cycle at five submerged arc furnaces producing silicomanganese at Transalloys Table V
Typical furnace reline, tap-hole rebuild, and tap-hole repair timelines Furnace
Complete or partial furnace reline
Tap-hole rebuild
Tap-hole repair
1
8 years
4 years
2 years – only front SiC block replacement
3
12 years
6 years
3 years– only front SiC block replacement
5
12 years
6 years
6 months – only front mickey replacement
6
8 years
4 years
2 years - only front SiC block replacement
7
12 years
6 years
6 months – only front mickey replacement
A complete furnace reline involves a total excavation of the furnace and replacement of all refractories, including sidewall, hearth, and tap-hole. A partial reline requires only the replacement of the sacrificial carbon ramming on the furnace floor. This will also allow for the repair of the tap-hole from the inside of the furnace. Excavation of the furnace burden and subsequent visual inspection of the furnace lining will determine the extent of the rebuild or repair.
Tap-hole rebuild Figure 11 shows the steps in the rebuilding of a SiC tap-hole. A rebuild is completed in four days (from furnace switch-out to first tap), including the warming up of the tap-hole refractories. Before switching out the furnace for a tap-hole rebuild the furnace needs to be melted down as far as possible, which will reduce the amount of loose burden that has to be removed from the tap-hole. The sequence of events after the last tap has been made and the furnace is switched out is as follows: cooling
down for 12 hours, removal of tapping launder, installation of a working platform in front of the tap-hole, removal of tap-hole face plate to expose tap-hole refractories, breaking out of tap-hole refractories with a Brokk machine, and allowing the remaining molten slag and metal to drain from the furnace (see Figure 11a). During the breaking out of the refractory and removal of loose burden and solidified slag, water cooling is applied to increase the rate at which the tap-hole area is cooled down. Final breaking out of tap-hole bricks is done carefully with a jackhammer to expose undamaged tap-hole bricks to establish a clean surface to tie-in the new refractory bricks. Figures 11b, 11c, and 11d show the installations of the SiC –blocks, lintels, and SiC bricks in furnace. The bricks below the block are started from a solid foundation or good floor. After the rebuild of the tap-hole is completed, the faceplate and launder flange are replaced. A steel pipe is placed in the tap-hole and only plugged with a small piece of tap-hole clay at the hot face of the pipe to assist with the opening of the first tap after start-up. After the launder is replaced, the tap-hole bricks are extended into the launder and then rammed with carbon material. It is of utmost importance to check the alignment of the claygun and drill after a tap-hole rebuild. In order for the newly build tap-hole to cure and bake properly, the furnace is switched in in STAR mode to allow for the gradual heating-up of tap-hole refractories over a 24-hour period. The same methodology is followed with the rebuild of a graphite block tap-hole.
Tap-hole repair Under normal conditions, repairing a graphite tapblock entails only the replacement of the front sacrificial mickey. This can be done on an eight-hour planned shutdown. The launder sand and refractory materials in the launder are cleaned out with a jackhammer and shovel to expose the full mickey.
Figure 11—(a) Brokk machine breaking out tap-hole; (b) placing of first SiC block; (c) two SiC blocks with SiC lintels; (d) repaired face plate and tap-hole flange
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Figure 12—(a) Breaking out of front graphite mickey with jackhammer; (b) installing new mickey; (c) steel plate over mickey holding it in position; (d) repairing launder with carbon ramming material
The old mickey is then removed by breaking it out with a jackhammer (Figure 12a) and replaced with a new micky (see Figure 12b). Removal of the first mickey allows for inspection of the second mickey and if required it can also be replaced. After the mickey has been replaced, a 20 mm steel plate is installed over the mickey to keep it in position (see Figure 12c). Carbon ramming material is then used to line the launder up to the tap-hole again (see Figure 12d). A new mickey does not require warming-up before it can be used. The same methodology is used when replacing the cold face SiC block. The estimate time for only replacing the cold face SiCblock is 12 hours.
Start-up after maintenance and repair The commissioning of a furnace after a major reline normally takes 7 days. The megawatt input per hour is regulated to allow heat soaking of refractories and carbon ramming, to drive off volatiles and moisture, and to allow for expansion of the refractories. Because the furnaces are open furnaces the refractories cannot be heated with gas and the warming-up is carried out with electricity using the furnace transformers in STAR mode configuration until the first tap is made. Gradual heating-up of furnace after a long shutdown, more than three months, is also important to allow for slow expansion of the refractories. This type of warm-up is completed in three to five days.
Conclusions At Transalloys the furnace linings are of an insulating and semifreeze lining design. SiC blocks and graphite blocks are used. The Journal of the Southern African Institute of Mining and Metallurgy
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Procedures and systems have been established to ensure an optimum life-cycle of furnace linings, tap-holes, and auxiliary tapping equipment. All furnaces are equipped with clayguns and drills. Proper plugging of tap-holes and only selective O2 lancing has extended the tap-hole lives noticeable. Well established heating and warming-up schedules for newly build furnace linings and tap-holes allows for optimum drying, baking, and expansion of refractory material.
Acknowledgements This paper is published by permission of Transalloys and Mintek. The contributions of our colleagues are gratefully acknowledged.
References Basson J., Curr T.R., and Gericke, W.A. 2007. South Africa’s ferro alloys industry present status and future outlook. Proceedings of Infacon XI: Innovation in Ferroalloy Industry, New Delhi, India, 18-21 February 2007. Indian Ferro Alloy Producers Association. pp. 3–24. Bezemer, K. 1995. The silicomanganese production process at Transalloys, Proceedings of INFACON VII, Trondheim, Norway, 11-14 June 1995. Tveit, H., Tuset, J.K., and Page, I.G. (eds). Norwegian Ferroalloy Producers Research Organization (FFF). pp. 573–580. Gous, J., Zietsman, J., Steenkamp, J., and Sutherland, J. 2014. Excavation of a 48 MVA silicomanganese submerged-arc SiMn furnace in South Africa – Part I: Methodology and observations. Proceedings of the 5th International Symposium on High-Temperature Metallurgical Processing. The Metals, Minerals and Materials Society, Warrendale, PA. pp. 255–269. Steenkamp, J.D., Sutherland, J.J., Hayman, D.A., and Muller, J. 2016. Tap-hole lifecycle design criteria: a case study based on silicomanganese production. JOM, vol. 68, no. 6. pp. 1547–1555. u VOLUME 119
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TAILING STORAGE CONFERENCE 2019 Investing in a Sustainable Future 16–17 October 2019
Birchwood Hotel & OR Tambo Conference Centre Johannesburg, South Africa
BACKGROUND Engineers designing tailing storage facilities are faced with a number of new challenges resulting from the encroachment of both formal and informal housing projects, legislation pertaining to water usage and pollution control, shortages of water for processing, and the requirements for tailing dam closure. This has resulted in the introduction of new designs for construction of more stable dams, alternative deposition methods, the introduction of non-permeable linings, and the capping of dams to encourage rehabilitation and minimize dust pollution. The shortage of water in Southern Africa has necessitated changes in dam design to minimize water usage by either reducing the amount of water to the dam or increasing the amount of water recovered. Understandably, new legislation has been passed to regulate the construction and operation of tailing storage facilities. This knowledge resides with few specialists in the industry, and the operators on the mines are sometimes unaware of the consequences of these changes for their operations. In many cases the operations engineer has been misinformed and the need has arisen to get the parties together to discuss the implications of the changes. Reprocessing of existing tailings adds to the complexity of operating a tailing storage facility, and many new operators have little or no reference material to assist them when planning a retreatment project. ▶ 572
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Ultimately, the design must be focused on the future closure of the facility, and this has been further complicated by changes in the minimum environmental requirements. Industry has requested a tailing seminar for interested and affected parties to share ideas and solutions with their peers. We invite all operations, designers, technology providers, and legislators to get together for what could be a very informative and successful event.
WHO SHOULD ATTEND „ Senior and operational management of mines „ Engineers responsible for tailing facility management „ Regional and national officials from DoE, DMR, DWA, and „ „ „ „
DEA Companies and individuals offering tailings processing solutions Researchers Environmentalists and NGO’s Legal representatives from mining companies
For further information contact: Camielah Jardine The Journal of the Southern African •Institute of Mining Metallurgy 2107 Head of Conferencing • Saimm P O Box 61127,and marshalltown Tel: (011) 834-1273/7 • Fax: (011) 833-8156 or (011) 838-5923 E-mail: camielah@saimm.co.za • Website: http://www.saimm.co.za 2019/06/24 8:24 AM
Effect of reclaimed bauxite on andalusite-based refractory castables for tundish applications
u
E. Matinde1, and S.L. Msibi1,2 Affiliation: 1 School of Chemical and Metallurgical Engineering, University of the Witwatersrand, South Africa. 2 Scaw Metals Group, South Africa. Correspondence to: E. Matinde
Email:
elias.matinde@wits.ac.za
Dates:
Received: 5 Sep. 2018 Revised: 15 Mar. 2018 Accepted: 1 Apr. 2019 Published: June 2019
Synopsis The robust nature of alumina refractory materials presents opportunities for their in-process reclaiming and re-use in non-critical applications. Virgin refractories can be substituted with different quantities of reclaimed materials without compromising the quality of the refractory products. In this study, the effect of reclaimed bauxite on andalusite-containing refractory castables for tundish applications was investigated. The recycled bauxite was formulated in different proportions to replace virgin andalusite in the castables. Standard tests were conducted to evaluate the physical, physicochemical, and thermochemical properties of the formulated products. Static corrosion tests were conducted using tundish slag in order to simulate the actual operating conditions. The results show that the standard flow behaviour, open porosity, and bulk density of samples containing up to 22 mass% reclaimed bauxite were comparable to those of the reference castable. However, high-temperature properties such as static corrosion at 1400°C and hot modulus of rupture at 1500°C indicated that there is an upper constraint to the amount of substitution. The study demonstrated direct potential savings of up to 10% when the reclaimed alumina replacement ratio was increased to 22 mass%. Keywords andalusite, bauxite, recycling and re-use, refractory castables, apparent porosity, bulk density, static corrosion test.
How to cite:
Matinde, E. and Msibi, S.L. Effect of reclaimed bauxite on andalusite-based refractory castables for tundish applications. The Southern African Insitute of Mining and Metallurgy DOI ID: http://dx.doi.org/10.17159/24119717/311/2019 ORCiD ID: E. Matinde https://orchid.org/0000-00017899-3311
Introduction Alumina-based refractories can be classified into low-alumina, containing less than 50% Al2O3, and high-alumina or bauxite, containing between 50 and 90% Al2O3. High-alumina refractories have widespread applications in high-temperature metal extraction processes due to their outstanding thermochemical and thermomechanical properties and stability (Rebouillat and Rigaud, 2002; Frulli, 2016). The typical chemical composition of refractory-grade bauxite consists of 85–90 wt.% Al2O3, 5–10 wt.% SiO2, 1–2wt.% Fe2O3, 3–4 wt.% TiO2, 0.4–0.8 wt.% CaO + MgO, and 0.20–0.8 wt.% K2O + Na2O (Nyoka et al., 2013). The thermochemical and thermomechanical properties of bauxite-based refractories rely principally on the conversion of the contained silica and alumina to the highly stable corundum (α-Al2O3) and mullite (3Al2O3· 2SiO2) phases (Rebouillat and Rigaud, 2002; Hill and Sehnke, 2006; Garbers-Craig, 2008; Frulli, 2016). Due to the high costs and supply-chain constraints of bauxite, the high-alumina refractories are increasingly being replaced by the cheaper and more readily available andalusite (Rebouillat and Rigaud, 2002; Hill and Sehnke, 2006; Garbers-Craig, 2008; Nyoka et al., 2013; Frulli, 2016). Andalusite is an industrial mineral containing about 57–59.5 wt.% Al2O3, 38–40 wt.% SiO2, 0.6–0.8 wt.% Fe2O3, 0.15–0.25 wt.% TiO2, 0.1–0.25 wt.% CaO, 0.1–0.15 wt.% MgO, 0.1 wt.% Na2O, and 0.2– 0.35 wt.% K2O (Nyoka et al., 2013). Andalusite forms part of the trimorphous group with stoichiometry Al2SiO5. This group also contains sillimanite and kyanite. These minerals yield a mixture of silica (SiO2) glass phase and mullite (3Al2O3· 2SiO2 or Al2O3· SiO2) when heated in air in the temperatures range 1100–1480ºC (Equation [1]) (Rebouillat and Rigaud, 2002; Garbers-Craig, 2008; Abou-Sekkina et al., 2011; Frulli, 2016). [1] The mullite phases are thermodynamically stable at high temperatures, and possess good refractory properties such as low thermal expansion, low thermal conductivity, good chemical stability, and excellent thermomechanical stability (Abou-Sekkina et al., 2011; Frulli, 2016). Since andalusite transforms to mullite at relatively low temperatures with minor volume expansion, it has widely been used in both fired bricks and in unfired (unfired bricks, castables, plastic mixes) refractory materials for various applications (Rebouillat and Rigaud, 2002; Frulli, 2016).
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Effect of reclaimed bauxite on andalusite-based refractory castables for tundish applications Recycling of alumina-based refractory materials When the degree of in-service damage to alumina-based refractories is such that stable operation of high-temperature processes cannot be ensured, the materials are broken down and discarded (Hanagiri et al., 2008). However, the non-destructible nature of these refractories poses environmental challenges in the form of disposal space and particulate dust emissions. Nonetheless, the robust nature of alumina-based refractories presents opportunities as well, particularly for their in-process recyclability and re-use, as well as their re-use in other industry sectors (Ndlovu, Simate, and Matinde, 2017). In recent years there has been an increasing demand to recycle the spent refractories by incorporating them into a variety of refractory finished products (Bradley and Hutton 2014; Ndlovu, Simate, and Matinde, 2017). In fact, spent alumina-based refractories can easily be re-used in the production of conventional monolithic refractory products, and have long been regarded as fit-forpurpose grades in the substitution of virgin refractory raw materials (Hanagiri et al., 2008; Bradley and Hutton, 2014). To date, the in-process recycling of reclaimed refractory materials has focused on developing refractory products for low-temperature and non-critical applications (Mazzanti et al., 2010; Schutte, 2010; Bradley and Hutton, 2014). Schutte (2010) investigated the potential for recycling of spent chamotte (calcined aluminosilicate clay), andalusite, and magnesia spinel refractory materials for use as monolithic materials, and proposed that new product formulations containing up to 70–80 wt.% reclaimed material can be used in non-critical and lowtemperature applications. Mazzanti et al. (2010) proposed that spent alumina refractory materials can be reclaimed and used in producing medium to high alumina-based castables with little or no compromise on the product quality. Based on the principle of downward position, Hanagiri et al. (2008) developed a methodology to recycle processed refractory waste wherein the reprocessed materials are used to produce materials of lower quality than the original materials. In other studies, Kumar et al. (2014) developed a new class of conventional and lowcement castables using 40 wt.% ferrochrome slag and 45 wt.% calcined alumina. The findings from these studies have proved the suitability of the products made from the recycled materials. Based on the findings from these studies, the broad objective of this study is to investigate the potential utilization of reclaimed alumina refractories in the production of refractory castables for non-critical applications. The recycling of spent refractories for castable applications requires optimization of interrelated properties such as particle size distribution (PSD), porosity, and thermochemical and thermomechanical properties, and hence a fundamental knowledge of their thermophysical, thermochemical (corrosion/ erosion resistance), and thermomechanical behaviour (thermal, volume, and mechanical stability). Their degradation behaviour when contacted with process materials such as molten oxides, molten metals, and corrosive and abrasive gases commonly found in most high-temperature metal extraction processes is particularly important from a safety perspective (Kumar et al., 2015; Sadik, El-Amraniand, and Albizane, 2014). Standard tests are often conducted in order to determine the physical, physicochemical, thermomechanical, and thermochemical behaviour of these materials under simulated conditions of service. Since refractory castables consist of matrix and aggregate phases, the optimization of the PSD of the starting materials
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is one of the key requirements in the formulation of refractory castables. According to Fruhstorfer and Aneziris (2014), the central problem in designing unshaped refractory castables involves the simultaneous optimization of the interrelated properties of stability, flowability, and density. In particular, the PSDs, especially those of the coarse aggregates, play a critical role in obtaining the desired final properties, due to their effect on the final particle-packing density. The particle density packing, in turn, has a significant effect on physical properties such as apparent porosity and bulk density (Schutte, 2010; Fruhstorfer and Aneziris, 2014). Fruhstorfer and Aneziris (2014) proposed that the PSD of coarse refractories consists of fine particle fractions (<0.1 mm), medium grain sizes (between 0.1 and 1 mm), and coarse grains (>1 mm). Schutte (2010) also proposed that the coarse aggregate fractions (1–3 mm and 3–5 mm) are usually considered for substitution during the recycling process. The matrix components, which include calcium aluminate cement binder, fillers, and deflocculating additives, are often not changed to reduce the effect of recycled materials on the flowability, workability, early strength, and volume stability of the formulated castable (Schnabel et al., 2014). As highlighted earlier, the thermochemical and thermomechanical properties of refractory castables are critical from a safety perspective. Properties such as the hot modulus of rupture (HMOR) and corrosion/erosion resistance are some of the most important characteristics of refractory castables that determine their applicability in various processes. The HMOR is determined by a three-point bending test and is indicative of the strength of the bond in the refractory at high temperatures, and is thus correlated to resistance to structural degradation, chemical slag attack, and abrasion wear (Hancock and Cannon, 2000; ASTM C583; Kumar, Kumar-Singh, and Strivastava, 2013). In contrast, the corrosion of refractory castable involves phenomena of chemical dissolution of refractory phases as well as the precipitation of new crystalline phases, while abrasion wear involves mechanical wear by high-velocity gases or flowing liquid phases. In other words, refractories must not only resist high temperatures but must also resist both corrosion and abrasion attack by liquid oxides and process gases (Kumar et al., 2015; Sadik, El-Amraniand, and Albizane, 2014). This study investigated the recycling of spent high-alumina refractories reclaimed from an electric arc furnace (EAF) roof at a ferrous foundry in Johannesburg, South Africa. The initial working lining thickness of the roof bricks is usually 270 mm, and the residual thickness at the end of each campaign is around 135 mm or more. This translates to 50% or more of the original thickness being unused before the furnace roof is demolished (see Figure 1). In some cases, the remaining thickness of the bricks can be as high as 80% of the original. This implies that a significant portion of the brick remains unused but is still discarded, thereby presenting opportunities for sufficient quantities to be recycled. The foundry consumes over 2 500 t of fresh alumina refractory materials per year. A 10 t tundish is used in the final steel-refining operation. Currently, the tundish lining consists of a 200 mm permanent/safety lining based on virgin andalusite and a 50 mm MgO· SiO2-based working lining. To date, no initiatives have been undertaken to recycle the spent refractory products. In this study, the effect of reclaimed bauxite addition on andalusitebased safety lining was investigated (see Figure 2). The working lining is replenished by gunning after casting several The Journal of the Southern African Institute of Mining and Metallurgy
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Effect of reclaimed bauxite on andalusite-based refractory castables for tundish applications
Figure 1—Condition of the furnace roof (a) before and (b) after service, and (c) the average residual thickness of individual bricks (photographs by S.L. Msibi, taken on 5 December 2016)
Figure 2—Depiction of (a) freshly lined tundish, and the simulated tundish lining configuration with the (b) new and (c) diminished working lining
ladles (Figure 2). In this case, the tundish permanent lining is considered to be a semi-critical application, as the temperatures in the tundish hardly exceed 1565°C even when the thickness of the MgO·SiO2 hot face or working lining diminishes due to slagmetal attack. Figures 2a and 2b illustrate the freshly lined tundish castable and the configuration of the new working lining before slag-metal attack, respectively. Figure 2c shows the schematic representation of the diminished working lining after slag-metal attack. As shown in Figure 2, the proposed application of the formulated safety lining castable does not involve direct contact with the metal or slag unless the working lining is penetrated or diminishes completely. Nevertheless, the safety lining castable should be robust enough to resist penetration by liquid slag and steel in order to prevent breakouts should the working lining diminish to a critical thickness.
Experimental procedure Three sets of tests were conducted. The first set of tests was conducted to evaluate the recyclability of the furnace bricks. The second set focused on the formulation of castables based on the specifications of the as-received reference castable. The final tests focused on the physicochemical and thermochemical properties of the castables. The following sections describe the characterization and standard tests that were conducted in order to simulate the physical, chemical, and thermal properties, and the performance The Journal of the Southern African Institute of Mining and Metallurgy
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of the formulated product under simulated service conditions in the tundish.
Selection of reclaimed roof bricks for recycling In order to determine the recyclability, brick samples were randomly selected from the EAF roof at different locations, as shown in Figure 3. Brick (a) was reclaimed from the periphery of the roof. Brick (b) was reclaimed from the electrode pot area, while brick (c) was reclaimed from the off-gas chute. It is clear that the wear profiles of these bricks were not uniform across the entire furnace roof. The selected brick samples were then characterized using scanning electron microscopy (SEM) and X-ray diffraction (XRD) techniques in order to determine the phase compositions.
Formulation of castables containing reclaimed alumina The castables were formulated based on different proportions of reclaimed alumina (referred to here as grog) to virgin castable. Castable mixtures of 15 kg per sample were prepared from the crushed reclaimed bricks and virgin andalusite (see Table I). Only the < 1 mm, 1–3 mm, and 3–6 mm aggregates of the castable were considered for substitution. As shown in Figure 4, the particle size distribution of the fractions considered for substitution was comparable to that of the virgin gunning material provided by the refractory supplier. Since the rheological and thixotropic behaviour of refractory castables is primarily influenced by amount of sub-micrometre VOLUME 119
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Effect of reclaimed bauxite on andalusite-based refractory castables for tundish applications
Figure 3—Selection and visual conditions of the reclaimed refractory bricks after campaign
Table I
Mixing ratios of formulation of castables (mass %) Description Binder premix Calcined alumina Bauxite grog 0–1 mm Bauxite grog 1–3 mm Bauxite grog 3–6 mm Andalusite 200 μm Andalusite < 1 mm Andalusite 1–3 mm Andalusite 3–6 mm Total (%)
Virgin sample (VS)
8 mass % grog (S1)
14 mass % grog (S2)
22 mass % grog (S3)
28 mass % grog (S4)
8 2 0 0 0 20 20 35 15 100
8 2 2 4 2 20 18 30 14 100
8 2 4 7 3 20 16 28 12 100
8 2 6 11 5 20 14 24 10 100
8 2 8 14 6 20 12 21 9 100
chemical composition of the formulated products was analysed by X-ray fluorescence (XRF) and the crystalline phases were analysed using XRD. These properties were then compared to those of the virgin andalusite sample.
Standard tests on formulated castables Castable flow behaviour and water addition
Figure 4—Particle size distribution of the as-received andalusite and formulations
particles and binders, a binding premix, consisting of calcium alumina cement (CAC), admixture, and microsilica was maintained at 8 mass% of the total formulation. The calcined alumina content also remained unchanged at 2 mass% based on the specifications of the reference castable from the refractory supplier. Furthermore, the 200 μm size fraction for andalusite was also fixed at 20 mass% of the formulation. No other additives, in the form of accelerator or retarders were added to modify the placing properties of the refractory castable. The current formulation was guided by the need to replicate the reference formulation, which does not contain additives. The
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The amount of water added was controlled to achieve the desired flow behaviour and workability based on the specifications of the reference castable. The flow behaviour was measured using the ASTM C1446 method, wherein a cone (100 mm base, 50 mm height, and 70 mm top diameter) was filled with the castable and vibrated for 30 seconds. The percentage self-flow (% S.F.) of the castable was calculated from the ratio of final (Df) to the initial (Di) diameter using the formula (ASTM 1446, 2011; Myhre, Sandberg, and Hundere, 1998): [2]
Bulk density and apparent porosity The bulk density and apparent porosity of formulated castables (50 x 50 x 50 mm, fired at 110, 600, 1000, and 1200ºC for 24 hours) were measured based on Archimedean evaluation method (ASTM C20-00; Harbison and Walker, 2005). The bulk density and apparent porosity values were calculated as: The Journal of the Southern African Institute of Mining and Metallurgy
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Effect of reclaimed bauxite on andalusite-based refractory castables for tundish applications [3] [4]
Static corrosion test The static corrosion tests were conducted in accordance with the DIN 51069 Blatt 2 test at 1400ºC for 12 hours in order to determine the potential for isothermal chemical interaction between tundish slag and formulated castables (DIN Blatt 2, 1972). The samples were moulded into cubes with a cylindrical section core- drilled to specified dimensions of 50 mm diameter and 35 mm depth. The cylindrical holes were then filled with an equal amount of slag and covered with a lid of the same refractory material. The refractory specimen was heated to 1400ºC in a carbolite furnace (heating rate 5–8ºC/min) and soaked at that temperature for 12 hours to allow for sufficient refractory-slag interaction. After firing, the samples were left to cool inside the furnace for 24 hours, and were then sectioned to evaluate the corrosion pattern of each brick using SEM analysis. Based on the initial dimensions (50 mm Φ and 35 mm depth), the corresponding volume change after reaction was calculated by measuring the change in the diameter and depth due to slag penetration. The volume change, estimated as percentage corrosion volume, was calculated according to Equation [5]. [5]
where Vi is the initial volume and Vf is the final volume.
Hot modulus of rupture HMOR tests were conducted at temperatures close to the actual operating temperatures in the tundish. Bar samples (25 × 25 × 150 mm) were heated to 1500ºC at a rate of 5–8ºC/min. The samples were soaked at this temperature, and then a three-point bending test was conducted until failure. The procedure was repeated six times. The HMOR values were calculated as: [6]
where P is the maximum force applied at rupture (N), L is the span between the supports (mm), b is the width of the specimen (mm), and d is the depth of the specimen.
Results and discussion Characterization of raw materials and reclaimed bricks Table II shows the composition of the as-received (unfired) andalusite and the reclaimed bauxite. Figures 5–7 show photographs, backscattered electron images, and the unnormalized chemical compositions of the three bricks. The three brick zones are labelled as: (1) the contaminated area (referred to as A 1-slag), (2) the interface (A 2 interface), and (3) the unreacted portion (A 3-brick). The compositional analyses of the respective portions also showed minimal slag/refractory brick interaction.
Characterization of brick (a) Figure 5 shows the characterisation of the brick excavated from the periphery of the roof. The preliminary visual inspection showed that brick (a) was intact, with minimum signs of thermal and chemical degradation. The A 1-slag zone consists predominantly of FeO, SiO2 and Al2O3, MgO, CaO, Na2O, and K2O, with a calculated basicity (CaO + MgO/SiO2 + Al2O3) of 0.34. The oxide phases can be attributed to result from the direct fly-off of process materials during introduction of powder materials into the EAF and/or the projection of liquid droplets from the bath at the impact points of the arc and/or of the oxygen jets on the steel bath (Guézennec et al., 2005; Brosnan, 2004). The presence of FeO tends to reduce the liquidus temperatures of the corrosion reaction products (Brosnan, 2004), and hence is likely to promote the early formation of liquid phases and an increased wear rate of the hot face of the brick (Lee and Zhang, 2004). Table II
omposition of as-received (unfired) andalusite and C reclaimed bauxite (mass%) Al2O3 SiO2 Fe2O3 TiO2 MgO + CaO K2O + N2O3 Andalusite 59.40 39.00 Recycled bauxite brick 81.80 12.60
0.85 0.15 1.60 3.20
0.20 0.30
0.40 0.50
Figure 5—Characterization of brick (a) showing (i) the pictorial representation of the brick, (ii) the backscattered electron image, and (iii) un-normalized EDS compositional (mass%) analysis The Journal of the Southern African Institute of Mining and Metallurgy
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Effect of reclaimed bauxite on andalusite-based refractory castables for tundish applications Alkali compounds (Na2O and K2O) are typical volatile compounds present in the electric arc furnace dust (de Buzin, Heck and Vilela. 2017; Sofilic´ et al., 2004; Prigent, Bouchetou, and Poirier, 2011). Similar to FeO, these materials also lower the liquidus temperatures of the deposited materials, and would thus contribute to accelerated wear of refractory bricks (Prigent, Bouchetou, and Poirier, 2011). Prigent, Bouchetou, and Poirier (2011) investigated the corrosion of alumina raww materials by sodium vapours. Their findings indicate corrosion phenomena characterized by a dissolution-precipitation process caused by Na2O-rich liquid phases. However, due to the low contents of Na2O and K2O in the materials deposited on the brick surface, the effect of these compounds on the liquidus temperature could not be quantified in the present study. The presence of TiO2 as an impurity in bauxite bricks (see Table II) is also detrimental to the integrity of the brick. TiO2 compounds tend to react with SiO2 to form low-melting glass phases when the bricks are fired and/or exposed to high temperatures in service (Hancock and Cannon, 2000; Harbison and Walker 2000).
Characterization of brick (b) The brick reclaimed from the electrode area exhibited signs of severe thermal and chemical degradation (see Figure 6). Refractory materials close to the electrodes tend to experience severe degradation due to the high temperatures from the arc and aggressive chemical and abrasive attack from the process gases, volatile metal compounds, and entrained materials. The un-normalized EDS results show that the slag components were rich in FeO (27 wt.%), SiO2 (17 wt.%) and Al2O3 (30 wt.%), Cr2O3 (17 wt.%), and NiO (8.3 wt.%). The presence of these oxides on the surface of the bricks can be attributed to the entrainment and oxidation of metal and slag droplets ejected from the steel bath during arcing (Sofilic´ et al., 2004). In particular, the high contents of Cr2O3 and NiO observed can be attributed to the highalloy steel scrap and alloying elements commonly used in the production of special grades of steel. It is clear that the severe thermal degradation could have increased the susceptibility of the brick to penetration by the process materials. The B2-interface also had uncharacteristically high amounts of copper and zinc oxides, which is indicative of the high level of contamination in the brick from the impurities present in the feed. Buzin et al. (2015) noted that electric arc
Figure 6—Characterization of brick (b) showing (i) the pictorial representation of the brick, (ii) the backscattered electron image, and (iii) un-normalized EDS compositional (mass%) analysis
furnace dust may contain varying levels of compounds of metals such as Pb, Cd, Ni, and Cu, which can contribute to the corrosion of the bricks.
Characterization of brick (c) Figure 7 shows the brick excavated from the off-gas chute area. Compared to brick (a), the SEM image of brick (c) also exhibited signs of crack formation (orange line) potentially resulting from the intermittent thermal stresses from the high temperature conditions experienced in the electrode area. Based on the preliminary observations highlighted in Figures 5–7, brick (a) was considered suitable for reclamation. Brick (b) was rejected on the grounds of having experienced severe thermal degradation and contamination by process materials. Although brick (c) showed fewer tendencies of degradation,
Figure 7—Characterization of brick (c) showing (i) the pictorial representation of the brick, (ii) the backscattered electron image, and (iii) un-normalized EDS compositional (mass%) analysis
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Effect of reclaimed bauxite on andalusite-based refractory castables for tundish applications it was rejected on the grounds of economic viability due to potentially low quantities of the reclaimable portions at the end of each campaign. After rejecting bricks (b) and (c) for recycling, further confirmatory tests were conducted on brick (a) and the findings were compared to the virgin bauxite brick in order to further determine the suitability of the brick for reclamation. The XRD analyses also confirmed insignificant differences in the mineralogical phases between the virgin and reclaimed bricks, as indicated in Table III. Based on these considerations, brick (a) was considered suitable for reclamation. Further formulation tests were then conducted on this sample. The findings in Tables II and III can be correlated to refractory phases for the bauxite bricks (> 80% Al2O3). According to Brosnan (2004), corundum and mullite are the main refractory phases present in fired (1450–1550ºC) bauxite materials. The amorphous phases shown in Table III are usually associated with the presence of glassy silica and other indigenous impurities (e.g. Na2O, K2O, Fe2O3, FeO, and TiO2) (Brosnan, 2004). However, the high content of these phases in the two samples is of substantial concern as they affect the solidus temperatures of bauxite refractories
Characterization of formulated castables As discussed earlier, the alumina-based castables were formulated based on the varying the ratio of reclaimed bauxite to virgin andalusite. Table IV shows the oxide composition of the formulated castables as measured by XRF. It is clear that the Al2O3 content of the castable increases with an increase in the amount of reclaimed bauxite in the castable. This is due to the fact that the reclaimed bauxite inherently contains a higher proportion of alumina compared to virgin andalusite (VS) (Harbison and Walker, 2000). Given the higher Al2O3 content of the formulated castable, the performance of the refractory castables can thus be assumed to approach that of bauxite. The CaO content of 1.5–1.8 wt.% for the formulated castable is also congruent with the CaO tolerance for low-cement castables (LCCs) as classified by other researchers (Parr et al., 1997; Lee et al., 2001; Garbers-Craig, 2008). The slight increase
Table III
Comparative quantitative XRD analyses of reclaimed and virgin bauxite bricks (%) Phases present
Reclaimed brick A
Corundum Mullite Magnesia Quartz Iron Amorphous
35.79 41.36 0.2 0.57 0.07 22.01
Virgin bauxite brick 38.68 32.69 0.25 0.04 0.11 27.88
in Fe2O3 can be attributed to the bauxite grog aggregate, which inherently picks up Fe2O3 as an impurity from process materials (Hancock, 2000; Pivinskii, Dyakin, and Perepelitsyn, 2015). Regrettably, the presence of impurities such as Fe2O3 tends to significantly affect the high-temperature strength properties of the formulated refractory castables.
Water addition and flow behaviour of formulated castables Table V highlights the flow behaviour of the formulated castable. The installation method for the tundish permanent lining requires that the flowability of the castable be acceptable immediately after mixing. Thus, the thixotropic, workability and predictable placement properties are particularly important for both vibratable and non-vibratable castables (Pivinskii, 1998; Schnabel, Buhr, and Dutton, 2012). In practice, a castable with good flow properties at a low water mixing ratio tends to form a homogenous and well consolidated product (Zawrah and Khalil, 2001). From Table V, it is evident that the flow properties of the grog-containing castables deteriorated with increasing addition of grog. The observed trend can be correlated to the increased water requirements of the higher grog formulations. Technically, the higher the amount of grog, the higher the amount of water required to maintain the same level of flow and workability properties (Hanagiri et al., 2008). In practice, poor flow and placing properties in castables are often solved by increased water additions, albeit at the expense of other critical parameters such as open porosity, bulk density, and strength (Hanagiri et al., 2008; Schnabel, Buhr, and Dutton, 2012). Zawrah and Khalil (2001) proposed that excess water can reduce the strength and increase the shrinkage of the castable, while too little water can result in voids and poor consolidation.
Apparent porosity and bulk density Table VI shows the results of open porosity and bulk density tests conducted at different temperatures. As expected, the apparent porosity increased gradually with an increase in the amount of grog. Frulli (2016) stated that the increase in apparent porosity can be attributed to the increased amount of bauxite
Table V
ater addition and flow behaviour of formulated W products Physical property
VS
S1
S2
S3
S4
Water addition (%) % Flow behaviour (after 30 seconds)
6 165
7 150
7 155
7.5 145
8 140
Table VI
Open porosity and bulk density of formulated products
Table IV
Property
Oxide composition of formulated castables (%) Oxide
VS
S1
S2
S3
S4
Al2O3 SiO2 CaO MgO Fe2O3
61 63 65 69 70 37 34 32 28 27 1.6 1.5 1.7 1.6 1.8 0.1 0.0 0.0 0.0 0.0 0.7 0.9 1.0 1.1 1.2
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Open porosity (%)
VS
110ºC 2.3 600ºC 4.1 1000oC 5.1 1200ºC 5.0
Bulk density (kg/m3) 110ºC 2.54 600ºC 2.52 1000ºC 2.50 1200ºC 2.58 VOLUME 119
S1
S2
S3
2.8 4.8 5.2 5.1
2.5 4.8 5.2 5.3
2.7 4.0 4.9 4.8 5.2 5.5 5.5 5.5
2.54 2.52 2.48 2.49 2.49 2.48 2.52 2.49
2.46 2.47 2.48 2.48 2.45 2.47 2.43 2.48
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Effect of reclaimed bauxite on andalusite-based refractory castables for tundish applications grog addition in the castable, since the refractory materials based on bauxite or chamotte exhibit higher porosity than andalusitebased materials. The effect of the amount of grog is clearly observed in samples S3 and S4 fired at 1000ºC and 1200ºC. Since these castables had higher water requirements (see Table IV) than the reference castable, the higher porosity can be attributed to disappearance of water of crystallization and hydration, which created porosity in the castable body structure without any meaningful bonding taking place (Kumar et al., 2014). At fixed firing temperature of 1000ºC, it can be inferred that the pore volume remained almost unchanged from VS to S3, but increased in S4 due to the apparent effect of grog addition (Zawrah and Khalil, 2001). Nevertheless, since low-cement castables (LCCs) usually possess open or apparent porosity values not higher than 10% (Lee et al., 2001), the values obtained for the formulated castables were considered acceptable for the proposed applications. Despite the marginal changes at intermediate firing temperatures of 600ºC and 1000ºC, it is clear that the bulk density decreased with increasing grog addition, particularly at 110ºC and 1200ºC. The observed trend is contrary to what was expected, since the higher order formulations contained more alumina, which is generally denser. Nonetheless, the observations from the present study are congruent with those of previous studies. Hutton et al. (2009) proposed that reclaimed refractory materials tend to have lower bulk density and higher porosity values compared to virgin materials. The bulk density is indicative of good particle size packing and porosity, and would thus correlate to the wear resistance of the refractory castables (Harbison and Walker, 2005). The apparent or open porosity, on the other hand, is a measure of the effective open pore space in a castable. Apparent porosity is thus indicative of the bulk density and wear resistance, i.e., the higher the bulk density and the lower the apparent porosity, the higher the resistance to chemical attack, slag-metal penetration, and abrasion resistance. Higher bulk density values are desirable in refractory castables, as such properties significantly affect the performance. Lee et al. (2001) proposed that the higher the bulk density, the higher the hot strength, and hence the higher the abrasion and corrosion resistance under service conditions. As such, the observed decrease in the bulk density with an increase in the amount of grog is likely to affect such high-temperature properties of the castable. As a result, it is important to further optimize the amounts of grog, reactive calcined alumina, and microsilica addition to enhance the packing density, and the nature and amount of bonding phases in the formulated castables (Pivinskii, 1998; Zawrah and Khalil, 2001; Parr, Fryda, and Wöhrmeyer, 2013; Sadik, El-Amraniand, and Albizane, 2014; Tomba-Martinez et al., 2017).
Static corrosion tests on formulated products Table VII shows the chemical composition of the slag used in the static corrosion tests. Due to the mismatch between the high
Table VII
Normalized chemical composition of slag (wt.%) CaO 32
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SiO2 FeO MgO Al2O3 31
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16
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7
Al2O3 content of the castable formulations and the process slag, interaction between the slag and safety lining can be expected at higher temperatures should the working lining diminish to critical levels. Although the safety lining is not in direct contact with the slag unless the MgO· SiO2 working lining diminishes completely, the corrosion tests were nevertheless conducted as a precautionary measure. Figure 8 shows the results of the static corrosion tests conducted at 1400ºC for 12 hours. As was expected, a high corrodibility, evaluated based on the Al2O3 pick-up in the slag phase, of the castable by basic slag, was observed. For example, the Al2O3 content of the slag phase in VS, S1, S2, and S3 increased to around 27% compared to 7 wt.% in the original slag. However, the relative impact of Al2O3 content and physical properties such as bulk density and open porosity of formulated castables is not clear up to S3. S4, with 28 mass% grog addition, higher open porosity, and lower bulk density, experienced higher slag attack as evidenced by the elevated pick-up of Al2O3 by the slag. Furthermore, it can be inferred that the slagcastable interaction occurred via the slag penetrating through the available pathways (open pores, cracks, phase boundaries, etc.), with the degree of penetration being controlled by the bulk density and apparent porosity of the castables (Pivinskii, 1998; Zawrah and Khalil, 2001; Poirier et al., 2008; Braulio et al., 2011; Sako, Braulio, and Pandolfelli, 2012; Parr, Fryda, and Wöhrmeyer, 2013; Kumar et al., 2014; Sadik, El-Amraniand, and Albizane, 2014; Tomba-Martinez et al., 2017). Thus, the corrosion behaviour indicates the need to further optimize the bulk density and open porosity properties of the formulated castables. In principle, corrosion reactions should be viewed as attempts by the system to achieve compatibility by progressing towards equilibrium (Brosnan, 2004; Sako, Braulio, and Pandolfelli, 2012). Sako, Braulio, and Pandolfelli (2012) proposed that Al2O3 dissolution can be expected as a result of the interaction of CaO-rich slags and Al2O3-rich refractory aggregate. Based on the approximate stoichiometric ratios of the CaO and Al2O3 of 2:1 observed for the slag phases, the reaction between the Al2O3 in the castables and the CaO in slag can be expected to occur via the formation of calcium dialuminate (CaO·2Al2O3) intermediate phases. The formation of the CaO·2Al2O3 phases from the corrosion attack on alumina-based refractories by process slag is also supported by findings from other studies (Lee and Zheng, 2004; Melcher, Reiter, and Harmuth, 2007; Poirier et al., 2008). Lee and Zheng (2004) proposed that the dissolution of Al2O3 occurs indirectly with the formation of intermediate phases CaO· 2Al2O3 and CaO·6Al2O3 under static conditions, with the direct dissolution of Al2O3 under dynamic and more aggressive slag conditions. Melcher, Reiter, and Harmuth (2007) also proposed that the corrosion mechanism of alumina spinel castables by steelmaking slags (basicity 0.80 CaO/SiO2+Al2O3 < 1.54) involves the formation of CaO·2Al2O3 and CaO·6Al2O3 phases from the concomitant diffusion of reacting species (Al2O3 from the refractory surface and CaO from the slag) towards the slag/ refractory interface. The change in the volume of corroded portions was used to further interpret the extent of corrosion of the Al2O3-castable by basic tundish slags. Based on the initial castable dimensions (50 mm Φ and 35 mm depth), the corresponding volume change after reaction was calculated by measuring the changes in the diameter and depth as a result of slag penetration. Table VIII The Journal of the Southern African Institute of Mining and Metallurgy
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Effect of reclaimed bauxite on andalusite-based refractory castables for tundish applications
Figure 8—Results for static corrosion tests for the formulated castables showing (i) the pictorial representation of the sample, (ii) the backscattered electron image, and (iii) un-normalized EDS compositional (mass%) analysis The Journal of the Southern African Institute of Mining and Metallurgy
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Effect of reclaimed bauxite on andalusite-based refractory castables for tundish applications Table VIII
Corrosion volume (%) of the formulated samples Sample Initial volume Average) Average Final % ΔV (mm3) Φ (mm depth (mm) volume (mm3) Corrosion VS S1 S2 S3 S4
68 750 68 750 68 750 68 750 68 750
53.5 53.2 53.5 53.2 54.0
35.5 35.5 35.0 35.0 35.5
79 836.33 78 943.48 78 711.88 77 831.60 81 335.57
16.13 14.83 14.49 13.21 18.31
shows that lower corrosion volume changes (%) were observed for S1, S2, and S3 compared against the reference VS. In essence, the relatively lower the percentage ΔVCorrosion, the higher the resistance to slag attack and the better the performance of the castable when exposed to slag conditions in the tundish (Parr, Wöhrmeyer, and Steffes-Tun, 2003). Likewise, S4, with 28 mass% grog additions, a higher open porosity, and a lower bulk density, experienced higher % ΔVCorrosion. The better corrosion resistance of S1, S2, and S3 may be attributed to the substitution of andalusite aggregate by bauxite with a higher refractoriness (Harbison and Walker, 2000). Harbison and Walker (2000) proposed that the refractoriness of alumina-based refractories is a function of alumina content, due to the fact that the aggregate grog, which is 80% bauxite, is added to replace < 60% alumina virgin andalusite in the castable. However, the benefits of higher refractoriness of bauxite were negated in S4 due to negative synergistic effects of water addition, increased open porosity, and lower bulk density which, in essence, potentially increased the susceptibility to corrosion attack. XRD analyses were conducted on the unreacted portions of the samples after the static corrosion tests (Table IX). Basically, the amounts of mullite and corundum phases increased with increasing amount of grog. This is due to the higher Al2O3 contents of the grog aggregates compared to the reference andalusite-based castable (Ødegård et al., 1998). Furthermore, the ratios of Al2O3 and SiO2 in the formulated castables are close to the stoichiometric ratio for the formation of the mullite (3Al2O3· 2SiO2) phases during firing and under service conditions. In essence, the formation of mullite can be represented to occur according to Equation [1] (Rebouillat and Rigaud, 2002; Hill and Sehnke, 2006; Garbers-Craig, 2008; Frulli, 2016). The volume and distribution of the mullite and corundum phases is particularly critical to the performance of the formulated castables. Mullite, in particular, provides chemical stability to resist corrosion attack, and as such, the better corrosion resistance of S1, S2, and S3 can be attributed to the increased amounts of these phases in the castables (Zawrah and Khalil, 2001; Rebouillat and Rigaud, 2002; Garbers-Craig, 2008; Abou-Sekkina et al., 2011; Frulli, 2016; Tomba-Martinez et al., 2017). Although S4 was expected to have higher resistance to slag attack due to the greater amounts of mullite and corundum phases, the integrity of the formulation was negated by synergistic effects of poor physicochemical properties such as bulk density and open porosity (Kumar et al., 2014). The effect of volume distribution of amorphous phases on the corrosion resistance of the castables is not clear from the findings. However, it can be inferred that the impurities in the raw materials (e.g., Fe2O3, Na2O, TiO2, and K2O in andalusite VS, and Fe2O3 in recycled grog) could have reacted with the Al2O3
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and SiO2 to form amorphous glass phases. Nevertheless, with the exception of S1 and S3, the amount of amorphous phases is congruent with the 10–15% proposed for high-alumina refractories (Yuan et al., 2015).
Hot modulus of rupture (HMOR) at 1500°C The thermomechanical properties of castables as a function of grog addition are presented in Figure 9. In general, castables with a higher recycled component demonstrated better HMOR properties at 1500ºC compared to the reference specimen. For example, S1 and S3, with comparatively higher mullite contents (see Table VIII), achieved the highest HMOR values of 2.12 MPa and 2.68 MPa, respectively. Congruent with findings from other studies, the formation of mullite significantly increases the hot strength properties of refractory castables (Myhre, 1994; Parr et al., 1997; Zawrah and Khalil, 2001; Rebouillat and Rigaud, 2002; Kumar, Kumar-Singh, and Strivastava, 2013; Kumar et al., 2014; Frulli, 2016). In fact, the higher HMORs for S1 and S3 can be attributed to both the higher mullite content and lower amount of amorphous phases, as indicated in Table VIII. Due to the ease of formation of liquid phases at high temperatures, amorphous glass phases have a significant effect on the thermomechanical properties of Al2O3-based refractories (Yuan et al., 2015). In addition, the values of HMOR of LCCs available in the literature are several orders higher than those obtained in this study (Chakraborty, 2004; Kumar et al., 2014), and the discrepancy is worth investigating in the future studies.
Economic considerations As discussed earlier, the recycling of reclaimed bauxite refractory bricks results in significant cost savings by reducing reliance on virgin andalusite raw materials. The proposed recycling and re-use of these materials will result in reduced costs of landfill disposal, thereby reducing the overall cost of ownership of the
Table IX
hase analysis of unreacted portions of castables P afterstatic corrosion test Phases Corundum Mullite Andalusite Cristobalite Anorthite Amorphous
VS
S1
S2
S3
S4
2.7 40 38 0.15 3.7 15
7.6 52 32 0.44 4.7 3.1
11 49 20 0.06 4.2 16
13 54 18 0.11 4.8 9
18 50 11 0.02 4.7 12
Figure 9—Comparative hot modulus of rupture values at 1500°C The Journal of the Southern African Institute of Mining and Metallurgy
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Effect of reclaimed bauxite on andalusite-based refractory castables for tundish applications Table X
Cost reduction from substitution of virgin andalusite per kg of formulated product
Raw material
Binder premix Calcined alumina Bauxite grog 0–1 mm Bauxite grog 1–3 mm Bauxite grog 3–6 mm Andalusite 200 μm Andalusite 0–1 mm Andalusite 1–3 mm Andalusite 3–6 mm Total cost ($) Total savings (%)
Unit costs 19.22 21.56 3.43 3.43 3.43 10.13 8.53 7.85 10.6
VS
S1 (8% grog)
S3 (22% grog)
S4 (28% grog)
Total cost ($)
kg
Total cost ($)
kg
Total cost ($)
kg
Total cost ($)
kg
Total cost ($)
75 20 0 0 0 200 200 355 150
1 442 431 0 0 0 2 026 1 706 2 787 1 590 10 000 0
75 20 20 35.5 15 200 180 319.5 135
1 442 431 69 122 51 2 026 1 535 2 508 1 431 9614 4
75 20 40 71 30 200 160 284 120
1 442 431 137 244 103 2 026 1 365 2 229 1 272 9287 7
75 20 60 107 45 200 140 248 105
1 442 431 206 367 154 2 026 1 194 1 947 1 113 9013 10
75 20 80 142 60 200 120 213 90
1 442 431 274 487 206 2 026 1 024 1 672 954 8530 15
refractories. In fact, it is proposed that at least 20% savings can be achieved by recycling alone (Bradley and Hutton, 2010). This section highlights the direct economic benefits from the formulations proposed in this study based on the replacement ratio of virgin andalusite. Table X illustrates the hypothetical costs of raw materials and the economic model for the cost reduction of substituting virgin raw material castable with reclaimed bauxite. The unit costs were assumed to include only the processing and sorting costs for the bauxite grog. From the standard test results, it can be deduced that S1, S2, and S3 offered a better compromise between cost and quality of the formulated castable product. Increasing the replacement ratio from 8 mass% to 14 mass% grog increased the direct potential savings from 4% to 7%. Potential savings of up to 15% can be achieved by further increasing the grog replacement ratio to 28 mass%. Although S4 offers the highest cost reduction ratio, the results from the selected quality tests revealed that higher replacement ratios can compromise the quality of the refractory castables. In practice, higher cost reductions than those shown in Table X can be achieved if the costs of dumping space and other salient benefits such as goodwill for environmental stewardship were to be included in the cost calculations.
Conclusion The study provides a pragmatic approach to the recycling of reclaimed bauxite bricks from the EAF roof. The effect of reclaimed bauxite on the performance of andalusitebased castables for tundish applications is discussed. The performance of formulated products was benchmarked to that of a refractory castable formulated from virgin andalusite. The diagnostic analyses of the bricks reclaimed from the periphery of the roof confirmed insignificant mineralogical differences to virgin andalusite bricks, and hence these bricks offered the best opportunities for recycling. The physical, thermal, and thermochemical behaviour of the products indicate that formulations S1 (8 mass% grog), S2 (14 mass% grog), and S3 (22 mass% grog) offered the best compromise between the cost and quality of the castables. Neglecting indirect costs such as conveying and the cost of dumping space, the study demonstrated that direct potential savings of up to 10% can be achieved at a grog replacement ratio of 22 mass% without compromising the product quality. The Journal of the Southern African Institute of Mining and Metallurgy
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S2 (14 % grog)
kg
Acknowledgements The authors acknowledge assistance from the University of Witwatersrand, Johannesburg and Scaw Metals Group.
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Kell hydrometallurgical extraction of precious and base metals from flotation concentrates – Piloting, engineering, and implementation advances K.S. Liddell1, M.D. Adams1, L.A. Smith1, and B. Muller2 Affiliation: 1 Lifezone Ltd & KellTech Ltd, Mauritius. 2 The Simulus Group, Australia. Correspondence to: K.S. Liddell
Email:
keith@kellprocess.com
Dates:
Received: 23 Jan. 2019 Revised: 27 Mar. 2019 Accepted: 3 Apr. 2019 Published: June 2019
How to cite:
Liddell, K.S., Adams, M.D., Smith,L.A., and Muller, B Kell hydrometallurgical extraction of precious and base metals from flotation concentrates – Piloting, engineering, and implementation advances. The Southern African Insitute of Mining and Metallurgy DOI ID: http://dx.doi.org/10.17159/24119717/602/2019
Synopsis Kell is a hydrometallurgical treatment option for the recovery of platinum group metals (PGMs), gold, silver, and base metals from flotation concentrates to refined products on site. The process is cyanidefree, while eliminating emissions of sulphur dioxide, arsenic trioxide, and other toxic species often emitted by smelters and roasters. Kell has been tested on a range of concentrates, including PGMs (UG2, Merensky, Platreef, Great Dyke, Great Lakes, Lake Superior and others), refractory gold-silver, coppergold, and polymetallic concentrates. High extraction efficiencies (>95%) are achieved for value metals (Pt, Pd, Rh, Au, Ag, Ni, Cu, Co) and secondary metals (Sb, Zn, Pb, and others). Studies comparing Kell with smelter-refining show economic benefits: capex 18–33% of smelting; opex 51–66%; electricity consumption 13–46%; and environmental benefits: greenhouse gas emissions 57–61% of smelting, acidification 5–27%, human toxicity 37–62%, and freshwater ecotoxicity 43–55% of smelting. Synergies for Kell plants alongside smelter-refineries include utilization of recovered acid from sulphur abatement systems and excess capacity in precious metal refineries. Kell is unconstrained by concentrate grade and impurities, allowing the co-location of Kell plants for treatment of non-smeltable concentrates alongside existing smelters. Kell has been demonstrated in a nine-week integrated pilot campaign at 1:1000 scale, supporting a bankable feasibility study for a 110 000 t/a plant treating a UG2-Merensky concentrate at Pilanesberg Platinum Mine (PPM), owned by Sedibelo Platinum Ltd, which has along with South African Industrial Development Corporation (IDC), invested in Kell. Zimbabwean Mining Development Corporation (ZMDC) has signed a Memorandum of Agreement (MOA) with KellTech for a centralized PGM concentrate processing plant. Keywords Kell, platinum group metals, base metals, hydrometallurgy, recovery, refining, environment.
Introduction Metalliferous concentrates from the mining of ore deposits have for millennia been treated by pyrometallurgical processing methods such as smelting and roasting. In the last 125 years, hydrometallurgical processes have played an increasingly important role, with the now widespread use of pre-oxidation techniques and cyanide leaching in gold ore and concentrate processing, and sulphuric acid heap leaching of low-grade copper ores. Currently, smelting remains the dominant approach for treatment of concentrates (Habashi, 2009), which are typically transported by road or rail to the nearest port for shipment to smelters worldwide. This global movement of millions of tons of concentrates annually has a detrimental impact on local communities and the environment and represents a large shift of employment prospects and economic value across international borders. Ore grades have been steadily decreasing, while the ore mineralogy of available resources is becoming more complex – with polymetallic ores becoming more prevalent as new resources (Adams, 2016). These factors have had an impact on the smeltability of concentrates due to the higher flotation mass pulls required from lower grade ores, resulting in lower concentrate grades. The direct result of this approach is an increase in gangue minerals and penalty elements such as arsenic, antimony, bismuth, chromite, mercury, and magnesium oxide, further constraining concentrate flotation recoveries and smelter payabilities. Arsenic is a problem for the economic and environmentally responsible processing of copper-gold and polymetallic concentrates as environmental regulations become globally more restrictive. China currently imposes a cut-off of 0.5% As in imported copper concentrate, beyond which it is not accepted by the smelters (AQSIQ, 2006). An increasing number of copper-gold and refractory gold concentrates have significantly higher arsenic contents, for example, at Cananea in Mexico, Chelopech in Bulgaria, Northparkes in Australia, Gortdrum in Ireland, Tampakan in the Philippines,
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Kell hydrometallurgical extraction of precious and base metals from flotation concentrates Aljustrael and Neves-Corvo in the Iberian Belt in Portugal and Spain (Long, Peng, and Bradshaw, 2012; Lane et al., 2016). Similarly, chromite and magnesium oxide constraints on platinum group metal (PGM) concentrate smeltability impose restrictions on concentrate flotation recoveries and smelter payabilities in the PGM industry. In response to these constraints, Kell was initially conceived by one of the authors (KL) and subsequently further developed (Liddell et al., 2011; Liddell and Adams, 2012). Kell is unconstrained by concentrate grade and impurities, allowing the treatment of concentrate streams that may not be readily smeltable by primary smelters. An additional environmental concern is the emission of sulphur dioxide by smelters and roasters directly into the atmosphere in some regions; emission requirements in more regulated jurisdictions necessitate the construction and operation of expensive wet sulphuric acid plants, which then require a market for the sulphuric acid. Co-location of Kell plants alongside smelter-refineries offers several potential synergies, including utilization of recovered acid from sulphur abatement systems at smelters and use of excess capacity in precious metal refineries. The long-established cyanide leaching process for the extraction of gold and silver is also meeting increasing resistance from stakeholders, primarily due to the human and ecotoxicity of the cyanide reagent, exacerbated by recent tailings spills and wildlife death events (Greenwald and Bateman, 2016). Current industry status is well expressed by Habashi (2014): ‘A new era in pressure hydrometallurgy is marked when roasting-leaching-electrowinning in the zinc industry was replaced in the 1980s by direct pressure leaching of zinc sulfide concentrate. Since then, pressure leaching has received recognition as the technology of the future. Large autoclaves of unprecedented size have been manufactured and installed in a number of metallurgical plants for recovering copper and nickel from their concentrates or liberating gold from refractory ores’.
Kell as a step-change technology Kell is a patented technology comprising four core sequential steps (depicted in Figure 1), all of which are well proven, commonly used in the metallurgical industry, and provide high recoveries of base and precious metals: 1. Aqueous pressure oxidation (POX) in an acidic sulphate medium to dissolve the sulphides and remove the base metals while minimizing dissolution of the precious metals. LME A
Figure 1—Simplified Kell process flowsheet
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Grade copper cathode is produced by solvent extraction and electrowinning (SX-EW). Excess iron in solution is removed by precipitation with limestone. Nickel and cobalt products are tailored to suit owner, site, and market conditions for the specific project, and may include mixed sulphide precipitate (MSP), high-purity nickel and cobalt metals, salts, or precursors. Conditions are similar to those used in industrial practice treating refractory gold or nickel concentrates targeting complete sulphur oxidation (typically 170–230°C with excess oxygen overpressure). 2. Atmospheric leaching for removal of soluble iron, and recovery of other HCl-soluble metals such as silver and residual base metals by means of scavenging sorbents and precipitation to secondary products suitable for further refining on site or externally. 3. Heat treatment of the solids residue to condition the mineral phases, rendering the material amenable to subsequent leaching. Typically PGM concentrates are heat treated at 500–1000°C using syngas. 4. Atmospheric oxidative leaching of the PGMs and/or gold in chloride/chlorine medium in a similar manner to that typically used in PGM refineries (50–90°C; 3–6 M HCl), with metal recovery by ion exchange and precipitation. Various selected products can be made using standard techniques to suit owner, site, and market conditions for the specific project, and may include intermediate reductive mixed precipitates, individual metals, salts or precursors. The separate leaching stages for the precious and base metals are designed to keep their respective chloride and sulphate chemistries disengaged, allowing for decoupled optimization of leach parameters and materials of construction for the two main value streams. Both base and precious metals leaching systems in Kell show very rapid kinetics – over 90% leach extraction in less than 60 minutes as shown in Kell test work leach rate data, compared, for example, with 24 hours typical for conventional gold cyanidation. For KellGold applications, the rapid leaching results in short pressure-oxidation autoclave residence times (RTs) of approximately 1 hour (approx. 1% of typical bacterial oxidation RTs of 5–6 days; (Miller and Brown, 2016). KellGold leach tank RTs are approximately 1 hour (4% of typical carbonin-leach (CIL) RTs of about 24 hours). KellGold therefore offers proportionally lower lock-up of contained gold metal value as working capital. Kell PGM plant design metal inventory pipeline lock-up times are estimated to be about 6 days for Pt and Pd and 4 days for Ni, Co, and Cu. A benchmark for overall conventional smelter-refining pipeline time for contained value metals may be estimated from reported segmental revenues and in-process metal inventory values (excluding refined metal) at Impala Refining Services (IRS), of 55 days and 60 days in 2017 and 2016, respectively (Impala Platinum, 2017). Hence, Kell pipeline times are estimated at approximately 10% of PGM smelter-refining pipeline times, with proportionally lower lock-up of contained metal value as working capital, a potential value release of some R800 million in the example cited. The relative impacts of Kell pipeline time on metal inventory lock-up compared with conventional concentrate processing in the PGM and gold industries are illustrated in Figure 2. Because sulphides and base metals as well as other acidconsumers are removed prior to chlorination, the benign gangue mineral particles pass almost intact through the Kell chlorination The Journal of the Southern African Institute of Mining and Metallurgy
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Kell hydrometallurgical extraction of precious and base metals from flotation concentrates Simulus Engineers. The LCA compared Kell against traditional pyrometallurgical (smelting and refining) treatment of both highsulphur and low-sulphur sulphide concentrates. Environmental performance was assessed in terms of impact categories, including:
Figure 2—Comparison of metal inventory value lockups for Kell and conventional processing methods
circuit, particularly so for chromite. The PGM or gold pregnant leaching solution (PLS) is clean, with low reagent consumption and amenable to production of low-impurity end products on site, such as 99.95% Pt, Pd, and Au sponges.
Environmental metrics Kell has many environmental benefits over traditional technologies (Smith, Adams, and Liddell, 2019). This alleviates concerns with SO2 and CO2 emissions, arsenic control, solid waste storage, water contamination, and cyanide usage. ➤ SO2 in vent gases is very low because sulphur is quantitatively removed as sulphate in an environmentally responsible manner. ➤ Arsenic is quantitatively removed in the front-end leaching circuits and is stabilized in phases such as basic ferric arsenate and scorodite, a well-accepted route that is standard practice in the extractive metallurgical industry. ➤ No cyanide is used, avoiding the need for cyanide transportation, detoxification, management, monitoring, storage, and the risk of wildlife and human toxicity. ➤ Solid waste may be co-stored as a very small (typically 2–5%) component of current flotation tailings arisings, creating no additional storage construction or management requirements. Alternatively, the Kell residues can be dry-stacked. ➤ Water contamination risk is minimized as water is re-used in the process. ➤ Low electricity consumption – 54–87% less compared to smelting using a streamlined cradle-to-gate life cycle assessment boundary, depending mainly on factors such as concentrate mineralogy and onsite or offsite reagent production. This results in substantially lower CO2 emissions. Smith, Adams, and Liddell (2019) provide additional data on these topics. Samples of combined tailings from a 9-week pilot-plant campaign for Kell processing of concentrate from Pilanesberg Platinum Mine (PPM) were subjected to geochemical and water extraction analysis, showing that waste type classification would remain unchanged (non-potentially acid generating, Type 3), and metals mobility from the tailings was determined to be very limited. To further quantify the environmental impacts and benefits associated with the Kell process, a preliminary streamlined environmental life cycle assessment (LCA) study was carried out by Energetics, supported by SysCAD mass-energy balance modelling by The Journal of the Southern African Institute of Mining and Metallurgy
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➤ Greenhouse gas emissions into the atmosphere (expressed in kg CO2 equivalents) ➤ Depletion of non-renewable fossil fuels resources (expressed in MJ (net calorific value) ➤ Acidification (relative effect of total emissions of acidic gases, expressed in kg SO2 equivalents) ➤ Toxicity (consensus-based, chemical specific characterization factors in USEtoxTM toxicity assessment models, expressed in comparative toxic units, CTU). The outcomes of the LCA indicated that for each ton of concentrate processed, Kell (when compared against conventional smelting and refining process on an equal output-value basis) results in significantly lower environmental impacts, primarily due to the substantially lower electricity requirements of Kell. The LCA study determined that Kell provides measurable benefit in each of the performance indicators examined (Figure 3): ➤ Greenhouse gas emissions: 39–43% reduction ➤ Resource depletion (fossil fuels): 33–60% reduction ➤ Acidification: 73–95% reduction ➤ Human toxicity: 38–63% reduction ➤ Freshwater ecotoxicity: (45–57% reduction).
Economic metrics Kell typically recovers a higher percentage of value metals (approx. 94–99% for Pt, Pd, Rh, Au, Ni, Co, Cu) than smelterrefining or cyanide leaching; hence, comparative discounted cash flow financial analysis is more appropriate than consideration of costs in isolation. Kell recoveries are generally higher than in other processes because (i) the POX step effectively liberates the PGM and gold particles from the encapsulating base metal sulphide matrix; (ii) PGM and other mineralogy is conditioned in the heat treatment step, producing readily leachable phases; (iii) chlorination conditions can be well controlled because the base metal sulphides and other ionic species have been removed from the system; (iv) specification of refined end products is integrated into Kell, is flexible, and may be selected to suit marketing and site location requirements; (v) relaxation of concentrate specifications can result in increased flotation recoveries.
Figure 3—Comparison of environmental profiles of Kell and smelter-refining (after Smith, Adams, and Liddell, 2019) VOLUME 119
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Kell hydrometallurgical extraction of precious and base metals from flotation concentrates Economic modelling based on comparative mass-energy balances carried out by Simulus Engineers shows that Kell outperforms existing pyrometallurgical processing for both capital and operating costs, for two different concentrate types (low S – approx. 1%; high S – approx. 6%), as illustrated in Figure 4. These outcomes are based on Kell design data from scoping or feasibility studies, with firm quotes for major equipment items, as well as mass-energy balances for Kell and smelter-refining compiled by Simulus Engineers. Smelter capital cost estimates were derived from prefeasibility study (PFS) data for a PGM concentrate containing approximately 8% S at 140 kt/a. Base metal refinery (BMR) capital cost estimates were derived from published costs for a previously considered BMR refurbishment at Zimplats’ Selous complex (Mandizha, 2015). These results are consistent with those from an independent comparative scoping-level options study undertaken by an international engineering company in 2012, as well as those from an investment case study conducted by an international consulting metals and commodities analyst company in 2018. Kell capex benefits arise from: (i) modular design that cements capex as part of quoted fabrication cost, avoiding capital blowout due to construction delays; (ii) short leach residence times, resulting in a relatively small plant footprint and tankages. Kell plant designs are scalable and have ranged in capacity (including refining) from 50 000 to 2 million ounces of precious metals per annum. Power consumption in Kell is low compared to the power required to melt the entire concentrate feed, including gangue
Figure 4—Comparison of economic profiles of Kell and smelter-refining with (a) reagent production on site, and (b) bulk reagent purchase
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minerals and fluxes, in conventional smelting. Chlorine consumption is low because chlorine consumers are removed in prior Kell process steps. Hydrochloric acid is recovered and recycled and is typically a nil operating consumption. On a case-by-case basis recovery and re-use of sulphuric acid may also apply. Generally, limestone sources near the mine site are available for any neutralization requirements The chlorination leaching system provides fast leaching rates (Aylmore, 2016), although halogens are reactive with other ore minerals, especially sulphides. Hence, under typical chloride leaching conditions, reagent consumptions are high if the ore or concentrate contains significant amounts of sulphide minerals (Aylmore, 2016; Mpinga, 2015). Kell eliminates this problem of reaction with sulphides, base metals, and other species because these are removed in the POX and atmospheric leach stages, leaving a residue that does not consume chlorine to any significant extent. The PLS is consequently quite low in impurities, enabling the recovery of precious metals to relatively pure products without excessive competition from other metals. Reducing minerals such as sulphides are removed prior to chlorination leaching and therefore the metal-chloride complexes tend to remain stable in solution under Kell chlorination leach conditions.
PGM concentrates Planned expansions and developments of southern African PGM concentrate production from the Platreef present operators with a potentially high capital and operating cost for new smeltingrefining capacity compared with the cost of a Kell plant, as illuminated above. A potential major expansion at Mogalakwena has been considered, which would involve building a third concentrator and would increase palladium output by 270 000 ounces per annum and platinum by 250 000 ounces per annum, on top of the approximately 500 000 ounces per annum of each metal currently produced (Seccombe, 2018). Platinum Group Metals reports that the definitive feasibility study (DFS) for the Waterberg project is due for completion in early 2019; the 2016 PFS considered a large-scale 600 000 t/month mine producing 744 000 ounces of 3PGM+Au per annum (Platinum Group Metals, 2018). Ivanplats delivered a DFS in July 2017 for a mine producing 476 000 ounces of 3PGM+Au per annum (Ivanhoe Mines, 2018). The net additional production of some 2 million ounces of 3PGM+Au per annum from the Platreef deposit will require additional concentrate processing and refining capacity. Currently there is a lack of electricity supply availability and certainty in both South Africa and Zimbabwe; the low electricity consumption in Kell processing therefore presents a further advantage. Sulphur dioxide emissions abatement is an increasingly important factor. Kell converts sulphur-containing compounds to stable insoluble residues, eliminating gaseous emissions. Sulphur abatement systems are currently not required for smelters in Zimbabwe (Gwimbi, 2017) and are installed at only some South African PGM smelters (Jones, 2005). Anglo American Platinum has announced a new R1.576 billion smelter abatement project at its Polokwane smelter (Anglo American Platinum, 2018) to meet South African legislation changes regarding new minimum emissions standards for SO2 stack emissions (DEA, 2013). Total capital expenditure on SO2 abatement technology is R2.5 billion – equal to Anglo American Platinum’s total profit for the 2017 financial year (Anglo American Platinum, 2018). The Journal of the Southern African Institute of Mining and Metallurgy
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Kell hydrometallurgical extraction of precious and base metals from flotation concentrates Gold and polymetallic concentrates Some 38% of the top 20 producers’ gold production in 2012 was from refractory sources, with an additional 14% in concentrate and 48% by free milling or heap leach cyanidation (Adams, 2016). Therefore, about half of current gold production is likely to be amenable to KellGold processing. Gold production trends are increasingly towards refractory and polymetallic feeds, with decreasing ore grades and more restrictions on cyanide use, effluent emissions, and increasing intensity of environmental activism. Traditional refractory gold roasters have been taken offline for environmental reasons; for example, Barrick and Newmont Mining’s KCGM Gidji circulating fluid bed (CFB) roaster in Western Australia was permanently shut down in June 2015 to eliminate atmospheric stack emissions. The Gidji roaster had previously accounted for around 99% and 90% of Newmont’s total annual sulphur dioxide and mercury emissions, respectively (Newmont, 2015). Cyanide-free KellGold treatment of various ores and concentrates has shown potential for application to a range of feedstocks (Adams, Liddell, and Smith, 2015), and more detailed batch testing and comparative economic assessments have since been completed (Adams, Smith, and Liddell, 2019), summarized in the following section. High metal recoveries of value elements into solution have been achieved from a variety of concentrates from batch tests (>95% of Au, Ag, Cu, Ni, Co, Zn, Pb, and Sb). Kell presents a substantial value recovery advantage over conventional alkaline cyanide processing, which does not readily recover silver from argentojarosites and other phases, for example. Rapid leaching results in small-size equipment units with substantially lower gold inventory lock-up compared with bioleaching-CIL. Considering overall residence times for bacterial oxidation (Miller and Brown, 2016), in the majority of cases a high level of oxidation (5–6 days) is required, and cyanide leaching of bioleaching residues requires 24–48 hours’ residence time. As shown in Figure 2, Kell requires about 1% of the bioleaching-CIL residence time, and hence gold lock-up, pipeline time, and tankage volumes for KellGold are much lower than for bioleaching. In several case studies (Adams, Smith, and Liddell, 2019), these factors, along with the typically higher metal recoveries, constitute a positive business case for Kell processing of many copper-gold, refractory gold, and polymetallic concentrates. Several scoping engineering studies have now been carried out for KellGold applications. The example shown in Figure 5 of a double refractory copper-gold-silver concentrate aims to produce refined 99.99% Au and Cu products on site, along with a silver product. KellGold eliminates smelter emissions of SO2 and As2O3 and no cyanide management, detoxification, or monitoring costs would be incurred in this cyanide-free application. KellGold represents a potential step-change in the gold industry.
The main Kell leaching and heat treatment unit operations are employed sequentially to assess (i) ultimate overall recoveries into solution and (ii) stagewise amenability to processing. Very high extractions of value metals have been achieved from a wide variety of PGM concentrates (listed in Table I), including UG2, Merensky, Platreef, Great Lakes, Great Dyke. and blends, indicating the robustness of the technology. To date, all PGM concentrates tested have been amenable to Kell processing, as illustrated in Table I. More extensive optimization and lockedcycle test work on each unit operation in the integrated Kell process has also been completed for several PGM concentrates. According to Mpinga (2015), the complexity of PGM mineralogy makes for a complex metallurgical scenario where a priori prediction of feasible processes is difficult. This is certainly the case for PGM treatment technologies that do not address the mineralogical complexity. In Kell processing, however, the complex PGM mineral phases liberated in the POX stage are converted to simple leachable metallic phases in the heat treatment step, eliminating the complexity issue completely. Moreover, Kell removes sulphur, base metals, amphoterics, and iron in the leaching stages prior to chlorination, resulting in a clean chlorination leach solution. Leaching of metallic PGMs in the Cl2/HCl system is well-established chemistry, readily described by simple equilibrium potential-pH diagrams (Pourbaix, Van Muylder, and de Zoubov, 1959) and with rapid leach rates (Landsberg and Schaller, 1971). Several refractory gold and polymetallic concentrates have been subjected to KellGold amenability testing, as shown in Table II. Additional leaching and metals recovery test work has also been completed on several of these concentrates.
Integrated pilot-plant operation Several pilot-plant campaigns were carried out at commercial metallurgical testing laboratories between 2011 and 2016 to test Kell for the processing of PGM polymetallic concentrates for recovery of Pt, Pd, Au, Rh, Ni, Cu, and Co. The most comprehensive campaign was carried out at Simulus Laboratories in Perth, Australia in 2016 on concentrate from PPM, and ran for 9 weeks as a decoupled integrated circuit at approximately 1000 scale-up ratio (illustrated in Figure 6). An industry norm of about 8000–15 000 is typical for hydrometallurgical plants scaled up
Kell process test work Batch amenability test work Concentrate samples of 5–15 kg are subjected to a standard batch amenability test using bench-scale equipment. The work, including inductively coupled plasma–optical emission spectrometry (ICP-OES) and inductively coupled plasma–mass spectrometry (ICP-MS) assay suites, is mainly carried out at Simulus Laboratories in Perth, Australia, supported by specialist assays and mineralogical investigations from other providers. The Journal of the Southern African Institute of Mining and Metallurgy
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Figure 5—Simplified KellGold flow sheet for a double refractory copper-gold-silver concentrate VOLUME 119
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Kell hydrometallurgical extraction of precious and base metals from flotation concentrates Table I
Overall Kell recoveries of value elements into solution (%) from PGM feeds Concentrate Overall recoveries into solution (%) ID
Type
Pt
Pd
Rh
Au
Ni
Cu
Co
A B C D E F G H I J K L M N O P
UG2 UG2 Merensky UG2-Merensky Platreef Platreef Polymetallic Great Lakes Polymetallic Great Lakes UG2-Merensky Polymetallic North America Great Dyke Great Dyke Polymetallic Australia Platreef Platreef Platreef
99 99 99 99 99 98 97 99 99 95 99 99 99 98 97 99
97 98 98 98 98 99 99 99 98 99 98 98 99 99 93 98
93 96 97 96 96 97 95 N/A 90 N/A 95 89 N/A N/A 93 94
99 97 99 99 99 96 96 99 99 99 98 99 92 97 94 97
97 95 99 98 99 99 99 99 97 99 98 99 99 99 99 99
93 96 99 99 99 99 99 99 96 99 98 99 99 99 99 99
99 83 98 93 99 99 99 99 95 98 96 96 93 99 98 98
Mean
98
98
94
97
98
98
96
Table II
Overall Kell recoveries of value elements into solution (%) from gold feeds
Concentrate Overall recoveries into solution (%) ID
1 2 3 4 5
Type
Au
Ag
Zn
Pb
As
Cu
Co
Sb
High-grade carbonaceous polymetallic ore Refractory gold concentrate Refractory gold polymetallic concentrate Double refractory Cu-Au concentrate Refractory gold concentrate
91 96 98 98 98
95 N/A 97 98 98
99 N/A 100 N/A N/A
95 N/A 97 N/A N/A
98 N/A 100 N/A N/A
100 98 99 99 N/A
N/A 97 N/A N/A N/A
N/A N/A 95 N/A N/A
Mean
96 97 99 96 99 99 97 95
The overall Kell PPM test work programme included the following components:
Figure 6—Integrated Kell pilot plant in 2016, showing 100 L long-term materials of construction testing system in centre foreground (photo with permission of Simulus)
from pilot to eventual commercial scale (Adams et al., 2004). An informal survey of seven autoclave-based nickel hydrometallurgy projects that were piloted since 2000 and culminated in operating plants showed that six of the seven (Ravensthorpe, Ambatovy, Rio Tuba, Ramu, Taganito, Gordes) scaled up from pilot to full scale at scale-up ratios typically of >10 000 (Adams et al., 2004; Collins et al., 2005; Collins and Vardill, 2005; Valle et al., 2016; Wilkinson, 2006; Tsuchida, 2015; Yesil and Iplikcioglu, 2015). One operation (Goro) built and operated a demonstrationscale plant at a scale-up ratio of 1000. Therefore, the Kell PPM campaign was considered well within industry norms.
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➤ Batch and pilot plant tests were completed in 2013 and 2016 on bulk samples of PPM concentrate (UG2-Merensky blends, approximately 95 g/t and 75 g/t 3PGM+Au, respectively). ➤ Monthly variability test work was completed in 2014 on 10 monthly mine production sample batches (with four duplicates) using Kell design operating conditions. ➤ Variability test work was completed in 2015 on five UG2/ Merensky blended sample batches using Kell design operating conditions. A 50:50 UG2:Merensky blend was selected for the 2016 pilot plant. A summary of recovery of value from concentrate (an expression of metal extraction from concentrate using 2016 Q4 metal prices) through the various stages of the overall programme is shown in Table III. It is noteworthy that the head grade was sequentially lowered during the stages of the programme, demonstrating consistent value recovery on lowering head grade from 110 g/t 3PGM+Au in 2014 to 73 g/t 3PGM+Au in 2016. Smelter feed grades are typically higher, at 130–340 g/t 3PGM+Au (Crundwell, 2011). This demonstrates the potential for Kell to treat higher mass pull concentrates by removing the smelter feed constraints, resulting in substantially higher flotation recoveries.
Long-term testing of materials of construction A 100 L continuously-agitated sealed tank system under The Journal of the Southern African Institute of Mining and Metallurgy
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Kell hydrometallurgical extraction of precious and base metals from flotation concentrates Table III
Overall Kell recoveries of value elements through the stages of pilot and variability test work for a PGM concentrate Concentrate type Test type Head grade, g/t 3PGM+Au Recovery of value, 3PGM+Au, %
Pilot conc. blend ~50:50 UG2:MER
Blend variability
Pilot conc. blend ~80:20 UG2:MER
Monthly variability
Pilot Plant 2016, steady state 73 96
Batch (avg.) 2015 85 96
Pilot Plant 2013, steady state 94 96
Batch (avg.) 2014 110 97
chlorination conditions was operated continuously at Simulus Laboratories for 7 months to test design-specified materials of construction for the chlorination circuit under operating conditions (shown in centre foreground in Figure 6). No wear of the tank or agitator impeller was observed during the trial and the materials were therefore considered suitable for use. The agitator seal and drive also performed well. No emissions triggers were experienced during the 7 months of operation.
2nd Edition (AusIMM, 2012). Trade-off studies for site-specific criteria, such as on-site reagent production, are included where warranted. The design methodology is modular fabrication and supply. Most of the process equipment, piping, valving, and instrumentation is pre-installed on skids or modules for transport to site. A small number of larger equipment items are installed directly on site.
Environmental testing
Engineering design status
Final neutralized Kell tails may either be dry-stacked or comingled with flotation tails and stored in the tailings storage facility (TSF). This allows for: (i) utilization of acid-consuming components of the flotation tailings (net acid-consuming); (ii) elimination of the need for separate lined dams for storage, thereby decreasing cost and eliminating environmental issues. Combined tailings from the pilot plant campaign described above were subjected to geochemical and water extraction analysis by an independent service provider, concluding that waste type classification would remain unchanged (non-potentially acid generating, Type 3), and that metals mobility from the tailings was very limited (Smith, Adams, and Liddell, 2019).
Over 15 scoping and prefeasibility studies and preliminary economic assessments have been completed for Kell applications to PGM, gold, and polymetallic concentrates, all showing attractive returns, particularly when Kell replaces smelting/ refining offtake contracts. A DFS, supported by batch optimization and variability and pilot-scale test work, was completed in 2013 for installation of a 280 000 ounce per annum Kell plant at PPM’s mine site. This was extended and upgraded, supported by circuit and set-point optimizations, additional bench-scale locked-cycle test work, and more extensive integrated piloting, resulting in a bankable feasibility study completed in 2016 (the block flow sheet is depicted in Figure 7). The overall body of work includes specific throughput studies, targeted test work, and Monte Carlo sensitivity analyses. Engineering and process simulation has been completed at construction design level.
Kell process engineering Engineering studies at scoping, PFS, and DFS level have been carried out for a range of Kell applications by Simulus Engineers, an engineering company in Perth, Australia specializing in hydrometallurgy, supported by targeted test work conducted by Simulus Laboratories on the same premises. For a given project, preliminary process design criteria (PDC) are prepared based on the client’s product requirements, test work results, prior work, and Kell process experience. Mass and energy balance outputs are generated for each case under consideration (e.g. throughput, grade, or concentrate type), using inputs from the PDC. The model outputs are used to size mechanical equipment, electrical loads, and motor selections. The capital cost estimate is built up from supplier quotations for major equipment items. For scoping-level studies only, earthworks, civil, structural, electrical, instrumentation, and piping costs are factored from previous Kell process studies and feasibility level designs. Installation costs for mechanical equipment and each of the disciplines are calculated from an estimate of installation hours and local costs. The operating cost estimates are developed from first principles. Reagent consumption rates are taken from the mass balance outputs. Unit costs are based on local supplier quotes provided by the client and vendors during recent studies. Labour requirements are estimated based on employee costs for site labour as supplied by the mine site. Energy consumption is built up from the equipment list and from site power costs. The target accuracies for the capital and operating costs are aligned with the AusIMM standard as stated in their Monograph 27, Cost Estimation Handbook, The Journal of the Southern African Institute of Mining and Metallurgy
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Industrial applications of Kell unit operations Unit operations in the Kell flow sheet are applied industrially under similar conditions in the extractive metallurgical and other industries. This body of experience and knowledge informs the design and materials selection of Kell unit operations and de-risks the engineering accordingly. A few examples from the database are given below to illustrate this point.
Figure 7—Kell block flow diagram showing metals recovery and refining components VOLUME 119
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Kell hydrometallurgical extraction of precious and base metals from flotation concentrates Pressure oxidation and base metals recovery
Kell process implementation
Pressure hydrometallurgy is well established in the processing of nickel sulphides and laterites, refractory gold, and some copper concentrates (Habashi, 2014; Thomas and Pearson, 2016). Kell POX conditions are selected to be similar to industrially proven processes in gold and base metals processing, targeting full oxidation of sulphides to sulphate, thus avoiding any complications of elemental sulphur formation. Base metal recovery methods follow established practice at primary base metal hydrometallurgical plants. Therefore, the process and materials performance and operability of the sulphate circuit are well established.
Comparison of Kell with smelter-refining
Heat treatment Flash dryer technology is already well utilized in the PGM industry on wet concentrate to produce smelter feed (Van Manen, 2006). Kell kiln conditions are similar to those used in industrially applied processes in other sectors, such as: ➤ Rotary kiln–electric furnace (RKEF) plants for smelting of nickel laterites to ferronickel (Stober et al., 2008) ➤ Calcination of kaolin for subsequent alumina leaching (Turner et al., 1982) ➤ Direct reduced iron production with coal-based rotary kilns (Yong-He and Sheng-Hui, 1999) ➤ Ilmenite reduction in a rotary kiln to allow iron removal by leaching and TiO2 production (Grey and Reid, 1973) ➤ Vanadium salt-roast and iron rotary-kiln reduction at Highveld Steel (Rohrmann, 1985) ➤ Conversion of spodumene to leachable form by heat treatment (Peltosaari et al., 2016), and ➤ Cassiterite reduction under CO-CO2 atmosphere (Su et al., 2016; Zhang et al., 2016). Mineralogy in the Kell heat treatment step is benign because sulphur in the kiln feed has been removed, in contrast to direct roasting of sulphide concentrates where bulk sulphur dioxide is produced. The conditions are non-oxidizing and non-corrosive, and materials performance is well established.
Chlorination and precious metals recovery Chlorination leaching technology has been applied in all PGM refineries since the advent of modern primary platinum ore production in the mid-1900s (Crundwell, 2011). Chlorination was first used in the recovery of gold from its ores in the Plattner, Deetken, Mears, and Newbery-Vautin processes in the mid- to late 1800s, before being displaced by cyanidation, primarily due to the high reagent consumptions incurred with chlorine when treating ores containing high sulphide grades (Adams, 2016). In Kell processing, these sulphides and other reagent consumers are removed prior to chlorination, resulting in low chlorination reagent consumptions. Criticisms of Kell include, for example, the misapprehension that ‘Cl2 gas is highly corrosive requiring special materials of construction and therefore results in both high capital and operating costs if such technology is adopted’ (Mpinga, 2015). Kell plant design incorporates modern high-temperature and abrasion-resistant, low-cost dual-laminate composite plastic materials such as thermoplastic-lined fibre-reinforced plastic (FRP). These materials have been well proven for similar duties in other industries such as steel pickling and gas scrubbing (OxyChem, 2013) and in some PGM refineries (Dunn, 2018).
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Potential synergies in the co-location of Kell within an existing smelter complex have been described earlier in this paper. Consideration of new concentrate treatment capacity at a greenfield site is also useful. Smelting melts the whole concentrate to recover the PGMs, Au and base metals – 80–95% of smelter feed is gangue. In contrast, Kell dissolves the base metal sulphides and precious metals, leaving most of the gangue minerals in their original form. The high energy input required to melt the gangue minerals is dictated by their heat capacities and latent melting heats, and this is the main reason for the much higher energy consumption in smelting than in Kell processing. A schematic comparison of Kell with smelter-refining of PGM concentrates is shown in Figure 8, which illustrates the main differences, particularly as regards elimination of the pyrometallurgical steps when adopting Kell processing. Kell processing comparisons with smelter-refining of PGM concentrates in this work and elsewhere (Smith, Adams and Liddell, 2019) are listed below. Similar outcomes would likely apply to copper-gold or refractory gold concentrates. ➤ Capital costs: 18–33% of smelting and refining (depending on concentrate mineralogy; on-site reagent production or bulk purchase) ➤ Operating costs: 51–66% of smelting and refining (depending on concentrate mineralogy; on-site reagent production or bulk purchase) ➤ Electrical energy consumption: 13–46% of smelting (depending on concentrate mineralogy; on-site reagent production or bulk purchase) ➤ CO2 equivalent emissions: 57–61% of smelting ➤ Shorter time to metal payment: <2 weeks rather than >5 weeks ➤ Metal recoveries comparable to smelting (approx. 96% value recovery) ➤ Processes lower grade, higher impurity concentrates – allows for increased concentrator recoveries ➤ No practical constraints on smelter grade or impurities such as As2O3, Cr2O3, and MgO ➤ Allows for higher concentrator mass pulls to lower grade, higher impurity concentrates
Figure 8—Comparison of Kell and smelter-refining block flow diagrams showing metals recovery and refining components The Journal of the Southern African Institute of Mining and Metallurgy
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Kell hydrometallurgical extraction of precious and base metals from flotation concentrates ➤ May result in additional concentrator recovery compared with smelting.
Kell implementation initiatives Pilanesberg Platinum Mine PPM is the wholly-owned operating subsidiary of Sedibelo Platinum Mines (SPM), which has also invested directly in the Kell licence for the SADC region. The South African Industrial Development Corporation (IDC) has further invested directly in the South African region Kell licence. The local Batkatla-baKgafela community is the largest shareholder in SPM and benefits from employment and training opportunities, and regional and social development. The project (Sedibelo Platinum Mines, 2018) has current overall resources for some 50 years of mining and milling 220 kt/month Merensky ore and 65 kt/month UG2 ore, producing 15 000 ounces of 3PGM+Au per month plus Cu, Ni, and Co in concentrate. The comprehensive body of Kell test work supports a positive bankable feasibility study for a 110 000 t/a (280 000 ounces per annum) plant treating a UG2-Merensky concentrate at PPM, with the following features: ➤ Feed concentrate: UG2 plus Merensky blend ➤ Design/construction: 21 months ➤ Commissioning: 3 months ➤ Ramp-up:18 months ➤ Construction phase: about 300 jobs ➤ Operations phase: 144 jobs ➤ Refining to Pt and Pd metals and added value products on site for direct end-user sale, or intermediate products ➤ Kell tailings co-deposited with flotation tailings, meeting regulatory and permitting requirements as shown by environmental test work (Smith, Adams, and Liddell, 2019) ➤ Potential for toll treatment of third-party concentrates. A 280 000 ounces per annum Kell plant has been approved for construction at PPM, subject to financial closure and forward mine planning.
Zimbabwe platinum producers Additional advances have been made in Zimbabwe, the world’s third largest platinum-producing country. The Zimbabwean government, represented by the Zimbabwean Mining Development Corporation (ZMDC), signed a Memorandum of Agreement (MOA) with KellTech Ltd. for a centralized metal concentrate processing plant. KellTech is actively engaging with the Zimbabwe Platinum Producers Association (ZPPA), as well as individual producers and owners, towards crystallization of this initiative. Test work and scoping-level engineering studies have been carried out for several Kell applications for in-country processing of various Zimbabwean PGM and gold concentrates, showing high value recoveries (>95%), and comparatively low capital and operating costs.
Additional Kell implementation initiatives Further test work and engineering studies for Kell applications at several other producers and development projects are being progressed for a range of concentrates, including PGMs (Platreef, UG2, Merensky, Great Dyke and elsewhere), refractory gold and silver, copper-gold, and polymetallic. In addition, Kell philosophy is being applied to the treatment and recovery of lithium (from brines, clays, and hard rock concentrates) and other materials. The Journal of the Southern African Institute of Mining and Metallurgy
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Conclusion Kell is a potential step-change in the low-emissions, cyanidefree hydrometallurgical recovery of metals from concentrates, with particular application to primary PGM, refractory gold, copper-gold, and polymetallic concentrates. Capital and operating costs and other economic drivers such as payment pipelines and working capital metal lock-up are a fraction of those for smelterrefining or oxidation-cyanidation, as are environmental impacts such as CO2, SO2, and cyanide emissions. Kell shows high metal recoveries across a broad span of concentrates and has been demonstrated in a 9-week integrated pilot-plant campaign and a 7-month large pilot-scale materials testing regime. This has culminated in delivery of some 15 scoping and prefeasibility studies for various concentrates, and a bankable feasibility study for a 110 000 t/a plant treating a UG2-Merensky concentrate at Pilanesberg Platinum Mine, for which construction is approved subject to financial closure and forward mine planning. The Zimbabwean Mining Development Corporation (ZMDC) signed a Memorandum of Agreement (MOA) with KellTech for a centralized PGM concentrate processing plant, towards which KellTech is actively engaging with the Zimbabwe Platinum Producers Association (ZPPA) and the mining industry.
Acknowledgements Industry partners who have provided samples and sponsorship, as well as fruitful discussions, are gratefully acknowledged. Sedibelo Platinum Limited is particularly thanked for their substantive financial, technical, and logistical support for Kell, and for permission to utilize specific operational and smelter PFS data. The support and investment of the Industrial Development Corporation of South Africa is acknowledged. The authors appreciate the excellent technical and engineering effort of staff at The Simulus Group, particularly Simon Willis, Brett Lawson, Duncan Cumming, Mike Storey, and Clint Perkis. Energetics personnel are thanked for their comparative life cycle assessment work on Kell and smelter-refining. The commercial skills of KellTech’s Chris Showalter are also well recognized.
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Platinum Group Metals Ltd. 2018. Platinum Group Metals Ltd announces increased confidence in large scale, palladium dominant deposit with strong gold credit at Waterberg JV South Africa, Press Release, 25 October 2018. http://www. platinumgroupmetals.net/investor-relations/news/press-releases/pressreleases-details/2018/Platinum-Group-Metals-Ltd-Announces-IncreasedConfidence-in-Large-Scale-Palladium-Dominant-Deposit-with-Strong-GoldCredit-at-Waterberg-JV-South-Africa/default.aspx [accessed 25 October 2018]. OxyChem. 2013. Hydrochloric acid handbook, 06/2013. Occidental Chemical Corporation, Dallas, TX. 51 pp. Pourbaix, M.J.N., Van Muylder, J., and de Zoubov, N. 1959. Electrochemical properties of the platinum metals. Platinum Metals Review, vol. 3, no. 2, pp. 47–53; vol. 3, no. 3. pp. 100–106. Rohrmann, B. 1985. Vanadium in South Africa. Journal of the Southern African Institute of Mining and Metallurgy, vol. 85, no. 5, May 1985. pp. 141–150. Seccombe, A. 2018. Unlike peers, cash-flush Amplats is ready for growth again. Business Day, 24 July 2018. https://www.businesslive.co.za/bd/companies/ mining/2018-07-24-unlike-peers-cash-flush-amplats-is-ready-for-growthagain/ [accessed 25 October 2018]. Sedibelo Platinum Mines. 2018. Resources and Reserves. http://www. sedibeloplatinum.com/assets/reserves-resources [accessed 21 September 2018]. Smith, L.A., Adams, M.D., and Liddell, K.S. 2019. The Kell Process: Efficient, lowenergy metals recovery with low environmental impact. Sustainable Mining 2019 – Proceedings of the 6th International Conference of Environment and Social Responsibility in Mining, Santiago, Chile, 4–6 September 2019. In press. Su, Z., Zhang, Y., Liu, B., Zhou, Y., Jiang, T., and Li, G. 2016. Reduction behavior of SnO2 in the tin-bearing iron concentrates under CO–CO2 atmosphere. Part I: Effect of magnetite. Powder Technology, vol. 292. pp. 251–259. Thomas, K.G. and Pearson, M.S. 2016. Pressure oxidation overview. Gold Ore Processing: Project Development and Operations. 2nd edn. Adams, M.D. (ed.). Elsevier, Amsterdam. Chapter 21, pp. 341–358. Tsuchida, N. 2015. HPAL in past, present and future. Proceedings of ALTA 2015 Nickel-Copper-Cobalt Conference. ALTA Metallurgical Services, Melbourne. 8 pp. Valle, L., Benz, M., Chalkley, M., Collins, M., Dobson, T., Holmwood, R., Malevich, A., Tuffrey, N., and Vrolson, R. 2016. Completing the Ambatovy ramp-up: the road to successful financial completion. Proceedings of ALTA 2016 Nickel-CopperCobalt Conference. ALTA Metallurgical Services, Melbourne. 19 pp. Van Manen, P.K. 2006. Process description and optimization of the flash dryers at Polokwane Smelter. Proceedings of the International Platinum Conference ‘Platinum Surges Ahead’. Southern African Institute of Mining and Metallurgy, Johannesburg. pp. 253–258. Wilkinson, T. 2006. Engineering of Coral Bay (Rio Tuba) nickel project: second generation HPAL? Proceedings of ALTA 2006 Nickel-Cobalt Conference. ALTA Metallurgical Services, Melbourne. 24 pp. Yesil, M.M. and Iplikcioglu, A.S. 2015. Gördes nickel cobalt HPAL project in Turkey. Proceedings of ALTA 2015 Nickel-Copper-Cobalt Conference. ALTA Metallurgical Services, Melbourne. 15 pp. Yong-He, H. and Sheng-Hui, X. 1999. CRIMM - a new process for production of directly reduced iron with coal-based rotary kiln. ICARISM - International conference on alternative routes of iron and steelmaking, Perth, Australia, 15–17 September 1999. Misra, V.N. and Holmes, R.J. (eds.). Australasian Institute of Mining and Metallurgy, Melbourne. pp. 223–228. Zhang, Y., Su, Z., Liu, B., Zhou, Y., Jiang, T., and Li, G. 2016. Reduction behavior of SnO2 in the tin-bearing iron concentrates under CO–CO2 atmosphere. Part II: Effect of quartz. Powder Technology, vol. 291. pp. 337–343. u The Journal of the Southern African Institute of Mining and Metallurgy
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Release analysis of coal fines: Evolution of the methodology and critical issues involved L. Sahu1, S. Bhattacharya1, and S. Dey2 Affiliation: 1 Department of Fuel & Mineral Engineering, Indian Institute of Technology (Indian School of Mines), India. 2 Mineral Processing Division, CSIR-National Metallurgical Laboratory, India. Correspondence to: L. Sahu
Email:
lksmineral@gmail.com
Dates:
Synopsis An ideal procedure to determine the limits of flotation-based separation, as is the case with washability analysis for density separation, does not exist. Therefore, a procedure called release analysis, considered to be the counterpart of the washability process in density separation, was developed to evaluate the flotation characteristics of coal fines. Since flotation is a complex process it is necessary to know the maximum recoverable yield that could be achieved at a target ash level and what could be obtained by release analysis. Due to a number of limitations in the initial procedures, modifications were recommended by various researchers to generate an optimum separation curve. This paper presents a review of, and a discussion on, the chronological evolution of release analysis methodology over the decades. Notwithstanding the availability of a standardized release analysis procedure, such as BS 7530, the paper highlights the critical parameters involved in the procedures, such as coal type, feed characteristics, choice of reagents and dosage, operator bias, and design features of the flotation cell. Keywords Release analysis, ideal procedure, advanced flotation washability, impeller speed, aeration rate.
Received: 25 Jan. 2018 Revised: 8 Jan. 2019 Accepted: 9 Jan. 2019 Published: June 2019
How to cite:
Sahu, L., Bhattacharya, S., and Dey, S. Release analysis of coal fines: Evolution of the methodology and critical issues involved. The Southern African Insitute of Mining and Metallurgy DOI ID: http://dx.doi.org/10.17159/24119717/17/409/2019 ORCiD ID: L. Sahu https://orchid.org/0000-00024610-1910
Introduction Coal is a heterogeneous substance consisting of various combustible constituents (macerals) and associated mineral matter. The kind and abundance of the various macerals present in the coal, such as vitrinite, semivitrinite, exinite, and inertinite, determine the type of coal. As macerals are formed by biological degradation of different parts of the plant material, they have different properties based on their original composition. Coal cleaning processes are dependent on the differences in physical properties between the coal components (the macerals) and also between the macerals and the mineral matter. Thus the cleaning characteristics of the coal might vary from one type to the next. The cleaning processes influence the distribution and the concentration of macerals in the various size and density fractions of the cleaned product. Flotation is the most commonly used method to clean and upgrade coal fines, and is typically applied in the size range of less than 500 μm. Coal being heterogeneous in nature, its composite property is dictated by its petrographic composition determined from maceral analysis, i.e. relative percentage share of macerals on a volumetric basis. Since different components, both macroscopic and microscopic, of coal have different chemical compositions, they show different degrees of floatability (Liu, Holuszko, and Mastalerz, 2017). For any coal, the macroscopic components vitrain and clarain are generally more floatable than durain and fusain (Arnold, 1986; Aplan, 1993). Analysis of microlithotypes describes the entire composition of a particulate cross-section of coal. The microlithotypes can be categorized on the basis of the number of macerals present, such as mono-, bi-, or tri-maceralic. The macerals in different proportions (at least 5% of each maceral group) within a 50 µm circle form a microlithotype (ICCP, 1963, 1971; Taylor et al., 1998). At a microscopic level, the efficiency of the flotation can be gauged from the partitioning of the microlithotype. The proportion of macerals in the microlithotype affects the floatability of the particles. Inertinite-rich coal, for example, is difficult to float. The separation of vitrinite and inertinite appears to be rank-dependent. The efficiency of any mineral response, as well as any maceral response, to flotation is dependent upon the proportion of the macerals and minerals on the surface of the particles (Hower and Parekh, 1986). Clay minerals and quartz dominate the mineral matter that is contained in typical fine coal flotation feeds. Flotation is a complex physicochemical process in which a collector is used to impart or enhance the hydrophobicity of the contact surface of floatable particles and a frother is used to stabilize the bubbles by reducing the surface tension of the pulp so that stable bubbles can drag the floatable particles to
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Release analysis of coal fines: Evolution of the methodology and critical issues involved the surface of the flotation pulp. The efficiency of separation in flotation is directly proportional to the probability of bubbleparticle attachment. The probability of bubble-particle collision for coarser particles is higher than that for finer particles. Thus flotation depends upon the size of particles and distribution of different size fractions in the feed material. The rate of flotation also varies with the particle size. Those particles having a size class of 300 × 75 µm show significant response, while 500 × 300 µm and < 75 µm particles have a slow rate of flotation (Sukumar, 2004). Relative density (RD) of the particles is closely associated with the particle size in the assessment of flotation rate. The higher the RD of the particle the lower the flotation rate. As flotation is selective and probability-based, a number of factors affect the performance of the flotation cell. Agitation of the pulp results in dispersion of the reagents throughout the pulp and collisions between bubbles and particles. This depends on impeller design (both rotor and stator), impeller speed, cell volume, cell aspect ratio, and cell design. The size of the bubbles formed, which is a key factor in flotation performance, is influenced by the aeration rate. As cell design features such as impeller design, structure and dimension of the cell, and operating parameters such as impeller speed and aeration rate vary from cell to cell, the results obtained in different cells are not identical, even under the same experimental conditions (Gorainm Franzidis, and Manlapig, 1995a, 1995b, 1996; Tabosa, Runge, and Holtham, 2016; Anzoom, Bhattacharya, and Sahu, 2017). As a result, process optimization is difficult where the limits of performance are unknown. To evaluate the performance of flotation the optimum boundary must be known so that an attempt to achieve the best possible separation can be made. Release analysis is considered to be an ideal separation procedure, like washability in froth flotation, and is used as the reference for batch or continuous flotation. In effect, the ideal flotation response of coal could be characterized by release analysis.
by removing the time limit of froth collection. Controlled aeration rate and impeller speed were introduced in the modified version (Figure 2). The froths were fractionated by varying the impeller speed and aeration rate instead of time. Both procedures were compared and the result of the modified procedure was found to be identical to those of the original procedure. The new procedure was considered to be simpler than the original procedure and more operator-friendly. Hydrophobic and hydrophilic materials were separated in the first stage, which produced both strongly floatable and weakly floatable particles in the froth. The froths were floated and refloated three times, or more if necessary. In the final stage, a froth from the first stage was fractionated by increasing the impeller speed and aeration rate. The problem of water decantation could be eliminated to a great extent by this procedure. A set of data for the simplified release analysis of coal is reported in Table I and Figure 3. A comparison of Dell’s original method and simplified release analysis is presented in Figure 4. As a larger amount of the froth was obtained, even in the first fraction, with higher ash, a time limit was required for the froth collection so that the yield-ash locus points of the release curve could be extended to a low ash level. In order to achieve this target, Cavallaro and Deurbrouck (1965) varied the impeller
Earliest procedure of release analysis Dell (1953) developed the first release analysis procedure as the ‘counterpart in froth flotation to float and sink analysis in the gravity concentration of coal’. Dell’s release analysis was trialed using copper ore, and was based on a step-bystep repeated flotation of products. Generally, the collector and frother dosages were typical for the average plant. Flotation feed was fractionated according to the time of collection and the procedure was thus called ‘timed release analysis’. Four floats were collected over a range of time intervals. The first and second floats were mixed and refloated. The second and third floats were refloated separately. This procedure was repeated two more times to minimize the amount of gangue particles entrapped by bubbles and entrained in water coming off with the froth. Finally, four concentrates with different grades and four tailings were obtained. The flow diagram of Dell’s original release analysis is shown in Figure 1. A release curve was prepared using a series of concentrates and tailings. The release curve showed the optimum result compared to other flotation experiments. The main drawback of this procedure was the necessity of decantating the froth every time before reintroducing to the next cell operation, to prevent pulp overflow.
Figure 1—Flow diagram of Dell’s original release analysis (after Dell, 1953)
Modification of the procedure Dell (1964) modified the original procedure of release analysis
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Figure 2—Dell’s modified release analysis procedure (after Dell, 1964) The Journal of the Southern African Institute of Mining and Metallurgy
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Release analysis of coal fines: Evolution of the methodology and critical issues involved Table I
Results of simplified release analysis (based on the data of Dell, 1964) Sample fraction Froth-1 Froth-2 Froth-3 Froth-4 Tailing Total
Weight, g 69.2 70.9 27.9 6.2 75.0 249.2
Weight, %
Cumulative weight, %
27.8 27.8 28.4 56.2 11.2 67.4 2.5 69.9 30.1 100.0 100.00
Ash, %
Weight of ash per 100 units feed
Cumulative weight of ash per 100 units feed
Cumulative ash, %
7.7 9.2 13.4 39.1 75.0 27.7
2.1 2.6 1.5 1.0 22.6
2.1 4.7 6.2 7.2 29.8
7.7 8.4 9.3 10.3 27.7
In terms of yield [1] where Y = yield of clean coal floated. In terms of recovery: [2]
Figure 3—Release curve (after Dell, 1964)
Figure 4—Comparison of the old method and the new method of Dell’s release analysis (after Dell, 1964)
speed with time as per Dell’s original 1953 procedure. The rest of the conditions were kept identical, and only time was limited due to the highly floatable nature of the feed. Since poor reproducibility was obtained in the release analysis results, possibly due to operator bias and certain other factors, an alternative method for optimum separation was developed by Nicol et al. (1983).
Classification of coal by release analysis based floatability criterion Since ash content and yield alone could not be used to assess separation efficiency, Hu (1975) introduced a floatability criterion which is comprised of separation efficiency, also called the capacity criterion, and the intensity factor or rate constant. The floatability criterion (Ek) can be calculated as:
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where R = recovery of combustible material K = flotation rate constant E = separation matrix = [ E1.I E1.II ] E2.I E2.II E1.I = combustible material present in clean coal, also called recovery E 2.II = mineral matter remaining in rejects E 1.II = combustible material contained in rejects E 2.I = mineral matter misplaced to clean coal. The floatability criterion is related to the separation matrix (E), which is a function of the recovery of combustibles, rejection of mineral matter, and possible misplacement between the two. Thus the separation matrix depends upon flotation performance, which might vary between cell types because of differences in their design features (Gorain, Franzidis, and Manlapig, 1995a, 1995b, 1996; Tabosa, Runge, and Holtham, 2016; Anzoom, Bhattacharya, and Sahu, 2017). Therefore the floatability criterion appears to be cell-dependent. Hu, Jin, and Bodily (1987) investigated the floatability evaluation of fine coal and suggested that the coals could be categorized on the basis of their floatability criterion (Ek) value into the categories very easy, easy, medium, and difficult to clean (Table II). Ek value of different coal samples could be calculated as in Table III.
Tree analysis: an alternative method Nicol et al. (1983) introduced a feed-independent ideal flotation response procedure called ‘tree analysis’, where the primary separation was followed by a number of scavenger and cleaner stages (Figure 5). Each product of flotation was subjected to
Table II
Classification of coal on the basis of Ek value aquired by timed release analysis Class of floatability Very easy Easy Medium Difficult
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Floatability criterion, Ek value >50 30-50 15-30 <15
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Release analysis of coal fines: Evolution of the methodology and critical issues involved Table III
Example of a floatability test by timed release analysis Coal sample
Yield, %
Helper Sunnyside Beaver Creek Wellington Hiawatha
76.9 84.7 71.1 63.0 55.6
Feed
Ash, %
Criteria
Clean coal
R%
ΣE
ΣK
Ek
9.3 5.9 8.7 5.6 12.7
86.1 89.4 18.2 64.9 59.2
39.3 38.4 30.7 20.6 16.4
1.3156 0.4081 0.3849 0.1901 0.3583
90.97 54.01 42.49 24.50 22.25
19.1 10.9 18.9 8.3 18.0
Figure 6—Partition curves obtained for different cells (after Nicol, Smitham, and Hinkley, 1994)
must be contained in the standard procedure so that it can be considered a kind of ideal response of flotation as a reference to evaluate the batch process. Figure 5—Tree procedure (after Nicol et al., 1983)
repeated branching flotation. Thus the entrained and entrapped particles carried in the froth could be removed and the hydrophobic particles remaining in the tailing could be recovered by a scavenger step. In each stage, the collector was added according to requirements that are possibly dictated by feed characteristics. All the mechanical parameters were kept constant throughout the experiment.
Partition curve Considering the tree analysis as a cell-independent procedure, Nicol, Smitham, and Hinkley (1994) explored the concept of a partition coefficient to measure coal flotation efficiency (analogous to the partition coefficient determined for density separation). Ash levels were grouped by ash ranges, e.g. 0–5%, 6–10%, 10–15%, 15–50%, and 50–100% (Figure 6). The partition coefficients were calculated by the ratio of the mass percentage of the product recovered within the ash range to the mass percentage of the total feed within the same ash range. The tree curve of feed provided the mass of coal present in each ash range. The mass of clean coal in each ash range was obtained by tree analysis of froth collected from the cell. The performance of the different flotation cells was evaluated on the basis of partition curves obtained for the various types of cells, as shown in Figure 6.
Criteria for standard procedure Pratten, Bensley, and Nicol (1989) specified some criteria that
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➤ The data analysis procedure should have a set of loci (cumulative yield, cumulative ash) so that a curve made by these points can define all the results obtained by the batch flotation. If a procedure has one single locus, it can be referred to for that ash content product. For example, for a target concentrate having 62% yield at 15% ash, the analysis cannot refer to a yield at 10% ash. ➤ The procedure must follow the separation based on surface properties. ➤ Separation made through the defined procedure must describe the optimum separation. ➤ The locus position should not vary with reagent dosage. ➤ The procedure must be simple, repeatable, and reproducible. Pratten, Bensley, and Nicol (1989) tested the tree process and concluded that it was reproducible and reported a closely related yield-ash curve (Figure 7). It is important to ensure close results after variation in the parameters, otherwise it should not be acceptable as a procedure. The only curve was shifted towards the higher yield ash points when the collector dosage was increased (Figure 8), which was due to the non-floatable particles being recovered in the froth phase. They also concluded that quite close results could be obtained by tree analysis while using various types of collectors. Their comparative results showed that tree analysis is superior to batch flotation tests and release analysis. However, that superiority could be obtained only at low ash levels. At high ash levels, tree analysis was inferior to release analysis (Figure 9). It appears, therefore, that no single procedure could be recommended for reference as an ideal separation such as washability for density separation. Pratten, Bensley, and Nicol further studied the dependence of the performance of the The Journal of the Southern African Institute of Mining and Metallurgy
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Release analysis of coal fines: Evolution of the methodology and critical issues involved
Figure 7—Reproducibility of tree procedure (after Pratten, Bensley, and Nicol, 1989)
Figure 10—Variation of collector type in tree procedure (after Pratten, Bensley, and Nicol 1989)
Figure 11—Effect of variation of collector and frother dosage and frother type (after Forrest, Adel, and Yoon, 1994) Figure 8—Variation of collector dosage in tree procedure (after Pratten, Bensley, and Nicol (1989)
Any small variation in a release curve using different reagents at different dosage levels was found to be within experimental error due to the variation in reagent type (Figure 11). Thus a release ash curve could be used to characterize a coal sample. In Figure 12, characterization of the flotation performance of the ROM coal and processed coal shows the release analysis as a tool for indicating ultimate separation. The release curve of the ROM coal showed sharp elbows, indicating that the majority of the mineral matter present in the coal is well liberated and hence is readily removed by flotation. The release curve of processed coal has shallow elbows, indicating interlocked particles with a high proportion of mineral matter.
British Standard procedure (BS 7530)
Figure 9—Comparison of batch flotation, release analysis, and tree analysis (after Pratten, Bensley, and Nicol (1989)
tree procedure on collector type, and concluded that the tree curve obtained (Figure 10) using different collectors was closely located, indicating the procedure to be independent of collector type. However, that was not the case with release analysis. In order to verify the effect of the reagent dosage and type on Dell’s procedure, Forrest, Adel, and Yoon (1994) had found that release analysis was independent of collector and frother dosage. The Journal of the Southern African Institute of Mining and Metallurgy
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In 1994, Dell’s simplified release analysis was adopted by British Standards after modifying the operational parameters for the evaluation of the flotation characteristics of coal (BS 7530, 1994). Though it is customary to report speed in revolutions per minute, BS 7530 (1994) envisaged the impeller speed to be fixed at 6 m/s throughout the experiment and the aeration rate was varied in the final stage. The reagents prescribed by the procedure are n-dodecane [CH3(CH2)10CH3] as collector and methyl isobutyl carbinol (MIBC) as frother. The procedure comprises two stages, with the first stage following an identical operational condition in each fraction of froth collection. Seven froth factions are collected, with reagent added at the stipulated dosage for each fraction. The froths collected in the first stage are mixed and reintroduced to VOLUME 119
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Figure 12—Characterization of ROM and processed coal (after Forrest, Adel, and Yoon, 1994)
the cell. In the second stage, the reagents are added once at the beginning of the test, and the froths are fractionated by varying the aeration time. The concentrates are obtained in a series of increasing ash. Details of the procedure are presented in the form of a flow diagram in Figure 13. Brown and Hall (1999) provided the background work for developing BS 7530 (1994) in detail.
Comparison of procedures Despite the number of procedures (Dell, 1953, 1964; Cavallaro and Deurobrouck, 1965; Nicol et al., 1983, Nicol, and Hinkley, 1994; BS 7530 1994) that have been developed to determine the limit of the flotation-based separation, analogous to the washability analysis of the float-sink test, there appears to be no standard procedure that could precisely specify the maximum yield at any ash content. Keeping this point in mind, Brown and Hall (1999) attempted to evaluate the most preferable procedure among the methods developed earlier. They compared the results from the two most popular procedures, which are release analysis and tree analysis, using coals from three different countries – Australia, South Africa, and the UK. Release analysis (BS 7530) produced a significant result at the low-ash end (initial part) of the curve, whereas the tree analysis method indicated optimum yield in the intermediate range of ash content. They could not conclude a preference for one single procedure or method of flotation. Therefore, they advised the selection of any standard procedure on the basis of the information required.
Application of wash water reduces the misplacement of gangue material through entrapment and all floatable particles remaining in the tailings are recovered by the repeated scavenging action. This ensures that all floatable particles participate in the next and subsequent stages. After completion of the first stage, all the froths are combined and subjected to a cleaning step. In the cleaning operation a low aeration rate is provided initially before the froths are fractionated by increasing the aeration rate.
Column release procedure The results recorded from a flotation column were found to be superior to those found in traditional procedures, release and tree analysis. In the flotation column, selectivity is better as particles and bubbles move in opposite directions. Thus particle-bubble collision is maximized and the column height, with a greater froth height, facilitates the retention of the entrained and entrapped particles, with a longer travel time to the froth collection zone. On the basis of the separation principle of the column, McClintock, Walsh, and Rao (1995) used the column to carry out release analysis by a so-called ‘column release procedure’. In the first stage, the concentrate is generated by continuous feeding of pulp. The froth is collected and kept aside and the tailings are subjected to a scavenging operation, where the froth is collected as a second concentrate and the tailings are again fed to the column. This operation is repeated until the flotation is complete.
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Figure 13—British Standards release procedure (after BS 7530, 1994) The Journal of the Southern African Institute of Mining and Metallurgy
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Release analysis of coal fines: Evolution of the methodology and critical issues involved Reverse release analysis Notwithstanding the availability of a standardized procedure such as BS 7530 (1994), it was observed by Randolph (1997) that some non-floatable particles also reported to the froth stream in release analysis. In order to minimize the misplacement of the non-floatable particles, Randolph (1997) introduced an alternative technique whereby the tailings are collected in the cleaning stage of flotation. This was named ‘reverse release analysis’ (Figure 14). The operating conditions followed in the cleaner stage of this procedure are in reverse of the sequence prescribed for release analysis (Dell, 1964), keeping the first stage the same. In the first stage, the floatable and non-floatable particles are separated using a high rate of air flow and high impeller speed. The tailing from this stage, designated as Tail-1, contains the highest ash of all the fractions and is positioned as the last coordinate of the curve. In the second stage, the rougher concentrate is fractionated in terms of tailing. The froth is collected until flotation is complete and an ultimate tailing is collected at each step. A high rate of aeration and agitation is maintained initially and then the operating conditions are in decreasing order, whereby a series of tailings with a decreasing level of ash is obtained. Dell’s simplified release analysis was not considered to be acceptable at low ash and low yield levels. It generated a concentrate with high ash at a high yield due to the misplacement of non-floatable particles. Although the reverse release curve constituted the coordinates of mostly tailings and a single concentrate, the curve was shown to be superior to the simplified release curve, with the points extending towards the low-ash region (Figure 15).
frother were used in the first stage. The second stage did not need any further collector addition. The frother, at a lower dosage, was added to produce a constant froth height of 90 cm which ensured the discharge of froth to the launder. The flow diagram of the procedure, termed advanced flotation washability (AFW), is presented schematically in Figure 16. The optimum separation curve formed by AFW has been claimed to be superior to the curves from other traditional procedures presented earlier. Therefore, this procedure was believed to be preferable for any target ash level. As it was independent of operator input, with zero operator bias, the reproducibility of the procedure was found to be acceptable (Figure 17). It also showed superior performance to a multistage column flotation operation. The comparative
Figure 15—Reverse release analysis results (after Randolph, 1997)
Advanced flotation washability (AFW) Mohanty, Honakar, and Ho (1998) investigated an alternative technique to evaluate the flotation response of a feed pulp sample. They set up a mechanical Denver cell and a packed bed flotation column. In the first phase, the hydrophobic particles were separated from hydrophilic particles by repeated flotation of the froth collected in the previous step. The final froth after repeated cleaning was fed to the flotation column, where the feed pulp was separated into six concentrates in order of their floatability by varying the aeration rate from 1.5 to 2.5 l/min. Kerosene as a collector and Dowfroth M-150 polyglycol as a
Figure 16—Schematic flow diagram of AFW (after Mohanty, Honakar, and Ho, 1998)
Figure 14—Reverse release analysis (after Randolph, 1997) The Journal of the Southern African Institute of Mining and Metallurgy
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Figure 17—Reproducibility of AFW (after Mohanty, Honakar, and Ho, (1998) VOLUME 119
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Release analysis of coal fines: Evolution of the methodology and critical issues involved results obtained for the ideal separation procedures are reported in Figure 18. The AFW technique also performed well for the desulphurization of coal. The same trend was observed for sulphur reduction, where AFW is superior to the release analysis (Figure 18b).
Floatability index The floatability index (FI) is an important factor that can assist in assessing the upgrading potential of fine coal. Mohanty, Honaker, and Govindaraja (1999) developed an equation by taking the area between ideal separation and no separation. The graph in Figure 19a was plotted by taking the X-axis as (100 minus ash rejection %) and Y-axis as (100 minus combustible recovery %). A higher FI value indicates easy-to-clean characteristics, whereas a coal with a lower FI value has difficult cleaning characteristics. The floatability index is also applicable for assessing sulphur rejection (Figure. 19b), where the FI is calculated by taking (100 minus total sulphur rejection %) instead of (100 minus ash rejection %).
FIash =
rea of the triangle created by no separation curve - Area A under the ideal separation curve Area of the triangle created by no separation
The FI values of two samples of coal fines from different locations are shown in Figure 19c. Illinois coal, which is of in situ origin, showed a FI value of 0.77, whereas Jharia coal of drift origin had a lower FI value of 0.32. This indicated that the drift coal would be difficult to clean.
Figure 19—AFW with no separation (after Mohanty, Honaker, and Govindaraja, 1999)
Factors affecting flotation performance by following different procedures Mohanty et al. (1998) studied the effect of collector dosage on the performance of release analysis. Lower collector dosages of 0.20 kg/t and 0.75 kg/t showed the optimum separation over a range of dosages (Figure 20). Increasing the collector dosage resulted in the recovery of semi-hydrophobic particles in the froth, lowering the grade of the concentrate product. Thus the curve shifted towards the higher-ash side of the graph. This indicated the sensitivity of release analysis to the collector dosage. These results contradicted those of the earlier study by Forrest, Adel, and Yoon (1994), where the results of release analysis were reagent dosage-independent and any variation was taken to be due to experimental error. Since fixing the collector dosage to produce the optimum separation is the first priority before proceeding to flotation, it added an extra task, viz. optimizing the collector dosage. As had been earlier reported
Figure 20—Variation of combustible recovery with collector dosage (after Mohanty et al., 1998)
in the methodology and results by Pratten, Bensley, and Nicol (1989), the tree analysis carried out by Mohanty et al. (1998) showed less variation in results with changes in the collector dosage rate.
Figure 18—The superior performance of AFW (after Mohanty, Honakar, and Ho, 1998)
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Release analysis of coal fines: Evolution of the methodology and critical issues involved A comparison of tree analysis and release analysis showed that each procedure had individual optimum separation regions. Tree analysis would be preferable in the low-ash and -sulphur region, while the release curve defined optimum separation in the high-ash and -sulphur region (Figure 21) The pulp density of the feed to flotation also affects the optimum separation. The mass fraction reporting to the froth is related to the carrying capacity, i.e. maximum mass rate at which the solids can be delivered into the froth. When the initial feed solids content is high, the total bubble surface area may not be adequate to transport all the particles that have already reported to the froth zone. As a result, selective detachment of the weaker hydrophobic particles may occur, and such particles fall back into the collection zone. Re-attachment of the particles to bubbles occurs repeatedly by reflux action (Dobby and Finch, 1986). Selectivity due to detachment and reattachment improves the separation at high feed solids content. Mohanty et al. (1998) showed that an initial high solids content in flotation feed had a significant effect on the release analysis results. An increase in the initial feed solids content from 8% to 16% significantly improved the separation, with little further improvement when the solids content was raised to 24% (Figure 22a). When combustibles recovery is plotted against sulphur content (Figure 22b), it is observed that feed solids content also has a certain effect on flotation performance as measured by combustibles recovery. According to Bhattacharya et al. (2017), the release analysis results are sensitive to particle size and collector type. If a size fraction responds well to a reagent, that does not necessarily
mean that the same reagent will deliver good results on any other size fractions of the same coal. Bhattacharya et al. (2017) studied the effect of feed size by treating coarse to ultrafine fractions with different collectors. They found that n-dodecane was significantly more effective on ultrafine size fractions of < 75 µm (Figure 23), where most coal particles were found to be liberated. However for the particles of size greater than 100 µm a synthetic collector of NALCO chemicals played a significant role in floating both liberated and unliberated particles (Figure 24). The combined effect of two different types of collectors might improve the flotation performance of the entire feed material. Bhattacharya et al. (2017) suggested the use of a blend of collectors on the basis of the response of different size fractions to the respective collectors and the mass proportions of those size fractions. It is postulated that an improvement in the release analysis results and flotation can possibly be brought about by the synergy between a collector effective for unliberated and partially liberated coarser particles and a collector effective for finer liberated particles.
Evolution of the procedures The evolution of the different versions of release analysis procedures over the last seven decades, beginning in 1953, and their modifications along with their attributes are highlighted in Table IV. The first ideal flotation-based separation procedure (Dell, 1953) was developed to evaluate copper flotation, and was later extended to coal flotation. A major difficulty encountered in the experimental work was the repeated decantation of water,
Figure 21—Comparative performance of release analysis and tree analysis (after Mohanty et al., 1998b)
Figure 22—Variation in separation performance with solids concentration (after Mohanty et al., 1998b) The Journal of the Southern African Institute of Mining and Metallurgy
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Figure 23—Effect of reagent type on release curves for finer feeds (after Bhattacharya et al., 2017)
a mechanical cell was used for the release analysis procedure known as column release (McClintock, Walsh, and Rao, 1995), which is similar to the simplified release analysis of Dell (1964). Because of repeated scavenging of the tailings to recover all the floatable particles, the procedure is time-consuming. Randolph (1997) therefore proposed a reverse release analysis in which the tailings were collected by repeated cleaning of froths, minimizing the misplacement of gangue particles. Advanced flotation washability (AFW), proposed by Mohanty, Honakar, and Ho (1998), comprised the use of a mechanical cell followed by column flotation. In this procedure the final froth obtained by repeated cleaning in the mechanical cell is fractionated into six froths in the flotation column by varying the aeration rate. Since an automated packed-bed flotation column is used, the procedure appears to be independent of the operator. Hence reproducibility of results is high, though the operation is time-consuming.
Considerations The discussion presented so far, in particular the one around Table IV, indicates that there remain a number of considerations, summarized below, which make the selection of a release analysis procedure very challenging.
Figure 24—Effect of the collector on release curves for coarser feeds (after Bhattacharya et al., 2017)
which was partially resolved by varying operational parameters such as aeration rate and impeller speed in order to control the pulp volume generated in the previous stages. Thus a simplified release analysis partially resolved the decantation problem (Dell, 1964). In an attempt to completely solve the decantation problem, Cavallaro and Deurobrouck (1965) varied the impeller speed with time with the aeration rate set to maximum. In view of the experimental bias in all these three procedures, perceived or real, an alternative method, tree analysis (Nicol et al., 1983), was developed in which roughing was followed by branching through the cleaning and scavenging stages. This route cleans the material thoroughly and the release curve generated covers all the points on a graph starting from a low ash-yield coordinate to a high ash-yield coordinate, indicating optimum yield. Tree analysis particularly showed superiority at the low-ash end of the curve. Its limitation was, however, the large number of products that were generated. Hu (1975) did not suggest any new procedure to evaluate the flotation response of coal. Hu, Jin, and Bodily (1987), however, proposed a methodology, called floatability criterion, to classify coal for flotation purposes, based on release analysis data, in a four-point scale from very easy to difficult to clean. The decantation problem in the timedrelease analysis appears to have been completely solved by British Standards (BS 7530, 1994) by dividing the pulp into froths by varying the impeller speed and aeration rate with time instead of varying only the time. A flotation column instead of
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➤ Feed characteristics—Results of release analysis vary with feed characteristics, raw coal versus cleaned coal (Forrest, Adel, and Yoon, 1994), and with feed particle size. Different size fractions show different flotation responses. This is particularly true for drift coal (Bhattacharya et al., 2017). ➤ Pulp density—Higher pulp density results in better separation (Mohanty et al., 1998). Pulp density is usually affected by the feed size, in particular, the mass percentages of < 75 µm and < 50 µm particles. Finding, the optimum pulp density, whatever is to be the procedure of release analysis, is therefore not an easy task. ➤ Reagent type—Release analysis results have shown frother dependency (Forrest, Adel, and Yoon, 1994) and collector dependency for the entire feed of < 500 μm (Pratten, Bensley, and Nicol, 1989) and for the constituent size fractions of the same feed, say 500 × 100 µm, < 1 00 µm, < 75 µm (Bhattacharya et al., 2017). ➤ Reagent dosage—Release analysis results have also shown dependency on dosage, in particular, that of the collector (Pratten, Bensley, and Nicol, 1989; Forrest, Adel, and Yoon, 1994; Mohanty et al., 1998). ➤ Operating conditions—Different experimental conditions, such as aeration rate, impeller speed, flotation time, froth cleaning methodology, and scavenging operations, all affect the results, whether the method be release analysis (Dell, 1964; Cavallaro and Deurbrouck, 1965; BS 7530 1994), reverse release analysis (Randolph, 1997), tree analysis (Nicol et al., 1983), or AFW (Mohanty et al., 1998). ➤ Cell type—Flotation performance is dependent upon cell design features such as tank geometry, impeller design, structure, and associated operational features (Gorain, Franzidis, and Manlapig, 1995a, 1995b, 1996; Tabosa, Runge, and Holtham 2016; Anzoom, Bhattacharya, and Sahu, 2017).).
Effect of petrographic composition on flotation of coal All the investigations on release analysis referred to in this study, except those carried out by Bhattacharya et al. (2017) and in an The Journal of the Southern African Institute of Mining and Metallurgy
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Release analysis of coal fines: Evolution of the methodology and critical issues involved Table IV
Different ideal separation procedures developed over the years Procedure
Researchers
Attributes
Limitation
Observations
Original Release Dell (1953) First procedure for analysis/ Timed ideal separation Release analysis applicable for flotation
Decantation of water due to the collection of forth on the basis of time. In case of highly floatable material, the time interval may not control the froth collection and hence a large amount of material may be floated initially. That might create problems for subsequent operations.
Copper ore was used for the initial study. The procedure is time-consuming and difficult.
Simplified Dell (1964) release analysis
Flotation rate could not be controlled due to fractionation of froth by varying aeration rate and impeller speed instead of time. For highly floatable feeds all material may come in the first or second froth product with little remaining in tailing. As a result, the curve may shift towards the high-ash end.
Kerosene as a collector and pine oil (or ethoxy butane) as frother wer used without providing justification for using these reagents.
Decantation of water as in original procedure was eliminated, making the methodology simpler in operation
Modified version Cavallaro and Impeller speed was varied Aeration rate was fixed at the maximum. of timed release Deurbrouck (1965) with time in order to acquire analysis low ash-yield locus. Tree analysis Nicol et al. (1983)
Fractionation was followed by A large number of products were generated. a number of cleaning and That made, handling of these products and scavenging operations, leading operational procedure difficult. to minimization of entrapment and entrainment of hydrophilic particles and to the recovery of residual hydrophobic material remaining in tailing.
Maximum aeration rate produced large bubbles. That facilitated the entrapment of gangue particles. As a result, the curve was unable to extend towards the low-ash end. Combustible material remaining in the tailings could be recovered by repeated scavenging. Gangue particles could be removed from clean coal by cleaning and re-cleaning. Thus more coordinates were obtained to plot an ideal separation curve.
Floatability Hu, Jin, and Floatability criterion based on The value of floatability criterion is procedure- criterion Bodily (1987) release analysis indicates relative dependent, hence cell-dependent too. cleaning characteristics of coal.
It only categorizes the coal on a four-point scale of very easy to difficult to clean. without predicting the maximum possible yield at target ash.
British BS 7530 Both impeller speed and aeration Prescribed a common non-commercial standards (1994) rate were varied for a specific time collector, n-dodecane, which might not procedure period in the second stage. perform well for every type of coal.
n-dodecane as collector and MIBC as frother were used. The combination appears to perform well only on in-situ or fully liberated coal.
Column McClintock release (1995) procedure
Since the entire separation is in Difficulties are faced in a repeated the flotation column, the yield ash cleaning of froth. locus can be extended to the lower ash region
Use of wash water minimizes the entrapment of gangue particles, thereby improving the concentrate grade.
Randolph (1997)
Froths were cleaned repeatedly Problems were encountered in to collect tailings in each cleaning tailing removal and reintroduction level, by which misplacement of of froth to cell non-floatable material to froth was minimized
Since highly hydrophobic coal was selected for the test work, no collector was used.
AFW Mohanty, Honaker, and Ho (1998)
Product cleaned by mechanical cell was subjected to column flotation. Hence entrainment and- entrapment of non floatable particles could be reduced.
Kerosene was used as the collector and Dowfroth M-150, a polyglycol, as the frother. The results showed the superiority of AFW over other traditional procedures.
Reverse release analysis
A large mass of feed sample is required to generate clean froth which is fed to flotation column.
isolated case by Brown and Hall (1999), have used in situ coal as feed material. All the samples used in the investigations appear to be vitrinite-, and possibly liptinite-rich, coals. The effect of maceral composition and the degree of oxidation of the coal do not appear to have been investigated, though these factors might affect coal flotation and therefore the release analysis. It is important to note in this context that Arnold and Aplan (1989) studied the response of coal components, both petrographic and chemical, to flotation. They used contact angle measurement techniques to assess the hydrophobicity of coal macerals. Generally, the rank of coal is determined by the percentages of C and O and vitrinite reflectance. These three factors provide a reasonable estimate of the flotation characteristics of the coal. The components containing carbon are mostly hydrophobic in nature. The hydrophobicity can be estimated from the contact angle of the particle. The contact angle increases with an increase in carbon content up to 88%, and after that point it decreases, as is the case with anthracite The Journal of the Southern African Institute of Mining and Metallurgy
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coal (Horsley, 1951; Gutierrez-Rodriguez, Purcell, and Aplan, 1984). This was verified using the sessile drop and captive bubble technique for measuring the contact angle of coal particles (Arnold and Aplan, 1989). The contact angle increased with an increase in oxygen content up to 5%, at which point a contact angle of 60–70° was observed. Thereafter the contact angle decreased rapidly to zero as the oxygen content increased to 20–25%. The contact angle was at a maximum at a vitrinite reflectance of around 1.0. For high-volatile bituminous coal, the hydrophobicity of coal macerals thus follows the order liptinite > vitrinite > inertinite (Arnold and Aplan, 1986; Aplan, 1993). The nature of the hydrocarbon chain and oxygen functional group dictates the hydrophobicity of macerals. The oxygen functional group allows water to adsorb on the coal surface due to the formation of strong hydrogen bonding between the water molecule and oxygen functional group. For higher rank coals, an increase in aromaticity results in a reduction in hydrophobicity, while an increase in the aliphatic group VOLUME 119
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Release analysis of coal fines: Evolution of the methodology and critical issues involved improves the hydrophobicity. Collector adsorption changes the surface chemistry of coal and thus its hydrophobicity. Selective adsorption of collectors on macerals with specific chemical composition influences the flotation performance (Wang et al., 2017; Holuszko and Mastalerz, 2015). Thus maceral concentration and maceral chemistry play an important role in recovery of froth. Characteristics of collectors such as their chemical composition, selectivity, and reactivity with the coal surface determine the performance of coal flotation (Holuszko and Mastalerz, 2015). Collectors should be such that they are able to increase the aliphatic group with minimization of aromaticity after adsorption so that hydrophobicity can be increased.
Summary and conclusions The first release analysis, also known as timed release analysis, was developed by Dell (1953) using copper ore. Dell (1964) subsequently simplified the procedure and used coal fines. Since then a number of researchers have suggested various modifications to the procedure. A British Standard (BS 7530, 1994) was developed for this purpose. To overcome the inadequacies of the release analysis procedures, alternative methods such as tree analysis, reverse release analysis, column release procedure, and advanced flotation washability (AFW) have been developed by other researchers. All the procedures mentioned show some dependence on factors such as feed characteristics, pulp density, collector and frother type and dosage, and experimental conditions such as aeration rate, impeller speed, flotation time, froth cleaning methodology, scavenging operations, and on cell type due to design features such as tank geometry, impeller design and structure, and associated operational features. Practically every study has been carried out using a different type of cell. Nearly all the investigations have been carried out using in situ coal as feed material. All the samples used in the experimental work appear to be vitrinite-, and possibly liptiniterich, coals. The effect of maceral composition and the degree of oxidation of the coal do not appear to have been investigated, though these features might affect the coal flotation and therefore the release analysis. This is particularly true for the inertinite-rich coals, typically of drift origin, because the hydrophobicity of coal macerals follows the order: liptinite > vitrinite > inertinite.
References
Anzoom, S.J., Bhattacharya, S., and Sahu, L. 2017. Dependency of batch coal flotation on cell types. NexGen Technologies for Mining and Fuel industries. Allied Publishers, Delhi. Vol. 2. pp. 1489–1494. Aplan, F.F. 1993. Coal properties dictate coal flotation strategies. Mining Engineering, vol. 45. pp. 83–96. Arnold, B.J and Aplan, F.F. 1986. Coal froth flotation: the response of coal and mineral particles to reagent and circuit variations. Advances in Mineral Processing. Society of Mining Engineers. pp. 351. Arnold, B.J. and Aplan, F.F. 1989. The hydrophobicity of coal macerals. Fuel, vol. 68. pp. 651–658. Bhattacharya, S., Jyoti, D., Sahu, L., Dey, S., and Singh, H. 2017. Flotation of low volatile coking coal fines. Transactions of the Indian Institute of Metals, vol. 70, no. 2. pp. 421–432. doi: 10.1007/s12666-016-0996-3 Brown, D.W. and Hall, S.T. 1999. Froth flotation: Preparation of a laboratory standard. Fuel, vol. 78. pp. 1621–1629. BS 7530. 1994, Methods for froth flotation testing of hard coal. Part 2. Evaluation of flotation characteristics. British Standards Institution, London. Cavallaro, J.A. and Deurbrouck, A.W. 1965. Froth flotation washability data of various Appalachian coals using the timed release analysis technique. Report of Investigation no. 6652. US Bureau of Mines. Dell, C.C. 1953. Release analysis - A new tool for ore dressing research. Recent Developments in Mineral Dressing. Institute of Mining and Metallurgy, London. pp. 75–84.
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Dell, C.C. 1964. An improved release analysis procedure for determining coal washability. Journal of the Institute of Fuel, vol. 37. pp. 149–150. Dobby, G.S. and Finch, J.A 1986. Flotation column scale-up and modelling. CIM Bulletin, vol. 79, no. 889. pp. 89–96. Forrest, W.R., Adel, G.T., and Yoon, R.H. 1994. Characterizing coal flotation performance using release analysis. Coal Preparation, vol. 14. pp. 13–27. Gorain, B.K., Franzidis, J.P., and Manlapig, E.V. 1995a. Studies on impeller type, impeller speed and air flow rate in an industrial scale flotation cell - Part 1: Effect on bubble size distribution. Minerals Engineering, vol. 8, no. 6. pp. 615–635. Gorain, B.K., Franzidis, J.P., and Manlapig, E.V. 1995b. Studies on impeller type, impeller speed and air flow rate in an industrial scale flotation cell - part 2: Effect on gas holdup. Minerals Engineering, vol. 8, no. 12. pp. 1557–1570. Gorain, B.K., Franzidis, J.P., and Manlapig, E.V. 1996. Studies on impeller type, impeller speed and air flow rate in an industrial scale flotation cell. Part 3: Effect on superficial gas velocity. Minerals Engineering, vol. 9, no. 6. pp. 639–654 Gutierrez-Rodriguez, J.A., Purcell, R.J. and Aplan, F.F. 1984. Estimating the hydrophobicity of coal. Colloids and Surfaces, vol. 12. pp. 1–25. Holuszko, M.E. and Mastalerz, M.D. 2015, Coal macerals chemistry and its implications for selectivity in coal floatability. International Journal of Coal Preparation and Utilization, vol. 35. pp. 99–110. Horsley, R.M. and Smith, H.G. 1951. Principles of coal flotation. Fuel, vol. 30. p. 54. Hower, J.C. and Parekh, B.K. 1986. Chemical/physical properties and marketing. Coal Preparation. Leonard, J.W. (ed.). 5th edn. AIME, Englewood, CO. pp. 16–19. Hu, W.B. 1975. Separation efficiency. Journal of Non-Ferrous Metals, vol. 6. pp. 40–51 Hu, W.B., Jin, R., and Bodily, D.M. 1987. Floatability evaluation of fine coal, Transactions of the AIME, vol. 282. pp. 1910–1915. International Committee for Coal Petrology (ICCP). 1963. International Handbook of Coal Petrography. 2nd edn. Centre National de la Recherche Scientifique and Academy of Sciences of the USSR, Paris/Moscow. International Committee for Coal Petrology, (ICCP). 1971. International Handbook of Coal Petrography. 1st Supplement to 2nd Edition. CNRS, Paris. Liu, J., Holuszko, M., and Mastalerz, M. 2017. Applications of micro-FTIR technique in studying hydrophobicity of coal. International Journal of Coal Geology, vol. 178. pp. 74–83 McClintock, W.W., Walsh, D.E., and Rao, P.D. 1995. Release analysis of a stratiform copper sulfide ore using column flotation. Minerals and Metallurgical Processing, vol. 12. pp. 112–117. Mohanty, M.K., Honaker, R.Q., and Ho, K. 1998. Coal flotation washability: Development of an advanced procedure. Coal Preparation, vol. 19, no. 1. pp. 51–67 Mohanty, M.K., Honaker, R.Q., and Govindaraja, B. 1999. Development of a characteristic flotation cleaning index for fine coal. International Journal of Mineral Processing, vol. 55, no 3. pp. 231–243. Mohanty, M.K., Honaker, R.Q., Patwardhan, A., and Ho, K. 1998. Coal flotation washability: An evaluation of the traditional procedures, Coal Preparation, vol. 19, no. 1–2. pp. 33–49. Nicol, S.K., Bensley, C.N., Teh, K.C., and Firth, B.A. 1983. The estimation of coal flotation response. Proceedings of the Congress for Improving Froth Flotation of Coal. Membrey, W. (ed.).Australian Coal Industry Research Laboratories. p. 116. Nicol, S.K., Smitham, J.B., and Hinkley, J.T. 1994. Measurement of coal flotation efficiency using the tree flotation technique. Proceedings. of the XII International Coal Preparation Congress, Krakow, Poland. pp. 741–746. Pratten, S.J., Bensley, C.N., and Nicol, S.K. 1989. An evaluation of the flotation response of coals. International Journal of Mineral Processing, vol. 27. pp. 243–262. Randolph, J.M. 1997. Characterizing flotation response: A theoretical and experimental comparison of techniques. M.S. thesis, Department of Mining and Minerals Engineering, Virginia Polytechnic Institute and State University, Blacksburg, VA. Sukumar, 2004. Size-density analysis of coal flotation process, M.Tech thesis, Indian School of Mines, Dhanbad, India Tabosa, E., Runge, K., and Holtham, P. 2016. The effect of cell hydrodynamics on flotation performance. International Journal of Mineral Processing, vol. 156. pp. 99–107. Taylor, G.H., Teichmuller, M., Davis, A., Diessel, C.F.K., Littke R., and Robert, P. 1998. Organic Petrology. Gebrüder Borntraeger, Berlin. 704 pp. Wang, W., Kumar, A., Holuszko, M.E., and Mastalerz, M.D. 2017. Selection of reagents based on surface chemistry as derived from microFTIR mapping of coal surface to facilitate selectivity in coal flotation. International Journal of Mineral Processing/Minerals Engineering - Special Issue after COM 2017. https://open.library.ubc.ca/cIRcle/collections/ facultyresearchandpublications/52383/items/1.0347213 u The Journal of the Southern African Institute of Mining and Metallurgy
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OBJECTIVES The conference will focus on improvement in health and safety, as well as the environmental impact the mining and metallurgy industries have on the local communities they serve while highlighting actions to be taken.
31 OCT - 1 NOV
Misty Hills Country Hotel & Conference Centre, Muldersdrift, Johannesburg
It will be a learning platform, allowing people to share ideas on health and safety, the environmental aspects which local communities’ need to be aware of, concerning the industry and its relationships with them. This conference aims to bring together management, DMR, Minerals Council South Africa, Unions, Health and Safety practitioners at all levels from the industry. Sharing best practice and successful strategies for zero harm and a value-based approach to health and safety. With the purpose of addressing the main challenges in the mining industry such as logistics, energy, employee safety, contractors and the communities at large. The conference should be of value to: ˙ Safety practitioners ˙ Mine management ˙ Mine health and safety officials The conferencemanagers should be of value to: ˙ Engineering ˙˙Safety practitioners Underground production supervisors ˙˙Mine management Surface production supervisors ˙˙Mine health andscientists safety officials Environmental ˙˙Engineering managers Minimizing of waste ˙˙Underground production supervisors Operations manager ˙˙Surface production supervisors Processing manager ˙˙Environmental scientists Contractors (mining) ˙˙Minimizing of waste Including mining consultants, ˙ Operations suppliers &manager manufacturers ˙˙Processing manager Education and training ˙˙Contractors (mining) Energy solving projects ˙˙Including mining consultants, Water solving projects ˙suppliers Unions & manufacturers ˙˙Education and training Academics and students ˙˙Energy DMR. solving projects ˙ Water solving projects ˙ Unions ˙ Academics and students ˙ DMR.
WHO SHOULD ATTEND
WHO SHOULD ATTEND
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SPONSORSHIP Sponsorship opportunities are available. Companies wishing to sponsor or exhibit should contact the Conference Coordinator.
For further information contact:
Head of Conferencing Camielah Jardine, SAIMM Tel: +27 11 834-1273/7 Fax: +27 11 833-8156 or +27 11 838-5923 E-mail: camielah@saimm.co.za Website: http://www.saimm.co.za
2019/06/24 8:24 AM
Diamonds – Source to Use — 2020 Conference Innovation and Technology
9 June 2020 — a technical and economic guide to diamond process engineering workshop and Technical Visits 10 June 2020 — Conference 11 June 2020 — Conference and Site Visit (SaB World of Beer) The Birchwood Hotel & OR Tambo Conference Centre, Johannesburg
BACKGROUND
T
he Diamonds – Source to Use conference series targets the full spectrum of the diamond pipeline, from exploration through to sales and marketing. The 2020 conference, the eighth in the series, will focus on advances in the mining and metallurgical aspects as well as many of the downstream and related industries
KEYNOTE SPEAKER L. Hockaday, Mintek – Renewable Energy Technology D. Collins Mac Consulting
OBJECTIVE The objective of the conference is to provide a forum for the dissemination of information relating to the latest tools and techniques applicable to all stages of the diamond industry, from exploration through mine design, processing, to cutting, marketing, and sales. WHO SHOULD ATTEND ➢ Geologists ➢ mineral (diamond) resource managers ➢ mining engineers ➢ Process engineers ➢ Consultants ➢ Suppliers ➢ Sales/marketing ➢ Diamantaires ➢ mine managers ➢ mining companies ➢ Students
Site visits ➢Epiroc South Africa ➢Multotec South Africa ➢SAB World of Beer Tour & Tasting
Conference Announcement June Journal.indb 608
TOPICS ➢ Geology and exploration ➢ mine expansion projects ➢ mining, metallurgy, and processing technology ➢ Rough diamond sales and marketing ➢ Cutting, polishing, and retail ➢ Synthetic diamonds ➢ Financial services and industry analysis ➢ industry governance, beneficiation, and legislation ➢ mine-specific case-studies 2 ECSA CPD points, 24 GSSA CPD points and 3 SACNASP CPD points will be allocated to all attending delegates
For further information contact: Camielah Jardine • Head of Conferencing · Saimm P O Box 61127 • marshalltown 2107 Tel: (011) 834-1273/7• Fax: (011) 833-8156 or (011) 838-5923 E-mail: camielah@saimm.co.za • Website: http://www.saimm.co.za
2019/06/24 8:24 AM
NATIONAL & INTERNATIONAL ACTIVITIES 2019 5–6 June 2019 — New Technology Conference and Trade Show ‘Embracing the Fourth Industrial Revolution in the Minerals Industry’ The Canvas Riversands, Fourways, Johannesburg Contact: Yolanda Ndimande Tel: +27 11 834-1273/7, Fax: +27 11 838-5923/833-8156 E-mail: yolanda@saimm.co.za, Website: http://www.saimm.co.za
Misty Hills Country Hotel & Conference Centre, Muldersdrift, 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 17–18 September 2019 — Hydrometallurgy Colloquium 2019 ‘Impurity removal in hydrometallurgy’ Glenhove Conference Centre, Melrose Estate, Johannesburg Contact: Yolanda Ndimande Tel: +27 11 834-1273/7, Fax: +27 11 838-5923/833-8156 E-mail: yolanda@saimm.co.za, Website: http://www.saimm.co.za
24–27 June 2019 — Ninth International Conference on Deep and High Stress Mining 2019 Misty Hills Country Hotel & Conference Centre, Muldersdrift, Johannesburg Contact: Camielah Jardine 16–17 October 2019 — Tailing Storage Conference 2019 Tel: +27 11 834-1273/7, Fax: +27 11 838-5923/833-8156 ‘Investing in a Sustainable Future’ E-mail: camielah@saimm.co.za, Website: http://www.saimm.co.za Birchwood Hotel & OR Tambo Conference Centre, Johannesburg Contact: Camielah Jardine 4–5 July 2019 — Smart Mining, Smart Environment, Tel: +27 11 834-1273/7, Fax: +27 11 838-5923/833-8156 Smart Society E-mail: camielah@saimm.co.za, Website: http://www.saimm.co.za ‘Implementing change now, for the mine of the future’ Accolades Boutique Venue, Midrand 31 October–1 November 2019 — International Mine Health Contact: Yolanda Ndimande and Safety Conference 2019 Tel: +27 11 834-1273/7, Fax: +27 11 838-5923/833-8156 Misty Hills Country Hotel & Conference Centre, Muldersdrift, E-mail: yolanda@saimm.co.za, Website: http://www.saimm.co.za Johannesburg 31 July–1 August 2019 — Entrepreneurship in the Minerals Industry Conference 2019 ‘Bringing ideas to life’ The Canvas Riversands, Fourways, 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 5–7 August 2019 — The Southern African Institute of Mining and Metallurgy in collaboration with the Zululand Branch is organising The Eleventh International Heavy Minerals Conference ‘Renewed focus on Process and Optimization’ The Vineyard, Cape Town Contact: Yolanda Ndimande Tel: +27 11 834-1273/7, Fax: +27 11 838-5923/833-8156 E-mail: yolanda@saimm.co.za, Website: http://www.saimm.co.za 18–21 August 2019 — Copper 2019 Vancouver Convention Centre, Canada Contact: Brigitte Farah Tel: +1 514-939-2710 (ext. 1329), E-mail: metsoc@cim.org Website: http://com.metsoc.org 19–22 August 2019 — Southern African Coal Processing Society 2019 Conference and Networking Opportunity Graceland Hotel Casino and Country Club, Secunda Contact: Johan de Korte Tel: 079 872-6403 E-mail: dekorte.johan@gmail.com Website: http://www.sacoalprep.co.za 28–30 August 2019 — IFAC MMM 2019 Symposium Stellenbosch Institute for Advanced Studies Conference Centre, Stellenbosch, Cape Town Email: info@ifacmmm2019.org Website: https://www.ifacmmm2019.org
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 13–15 November 2019 — XIX International Coal Preparation Congress & Expo 2019 New Delhi, India Contact: Coal Preparation Society of India Tel/Fax: +91-11-26136416, 4166 1820 E-mail: cpsidelhi. india@gmail.com, president@cpsi.org. inrksachdevO1@gmail.com, hi.sapru@monnetgroup.com 18-22 November 2019 — Coal Pillar Design Course The Estuary Hotel, South Coast, KwaZulu-Natal Contact: Yolanda Ndimande Tel: +27 11 834-1273/7, Fax: +27 11 838-5923/833-8156 E-mail: yolanda@saimm.co.za, Website: http://www.saimm.co.za 2020 25–26 February 2020 — SAMCODES Conference 2020 ‘Good Practice and Lessons Learnt Industry Reporting Standards’ The Birchwood Hotel & OR Tambo Conference Centre, 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 25–29 May 2020 — The 11th International Conference on Molten Slags, Fluxes and Salts The Westin Chosun Seoul Hotel, Seoul, Korea Tel: +82-2-565-3571, Email: secretary@molten2020.org http://www.molten2020.org 9–11 June 2020 — Diamonds – Source to Use — 2020 Conference ‘Innovation and Technology’ Birchwood Hotel & OR Tambo Conference Centre, 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
4–5 September 2019 — Surface Mining Masterclass 2019 Birchwood Hotel & OR Tambo Conference Centre, Johannesburg Contact: Camielah Jardine 18–22 October 2020 — IMPC XXX International Mineral Tel: +27 11 834-1273/7, Fax: +27 11 838-5923/833-8156 Processing Congress E-mail: camielah@saimm.co.za, Website: http://www.saimm.co.za Cape Town International Convention Centre, Cape Town 11–12 September 2019 — Revitalising exploration activity Contact: Camielah Jardine in southern Africa Tel: +27 11 834-1273/7, Fax: +27 11 838-5923/833-8156 ‘Potential for Exploration’ E-mail: camielah@saimm.co.za, Website: http://www.saimm.co.za
The Journal of the Southern African Institute of Mining and Metallurgy
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VOLUME 119
MAY 2019
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Company affiliates The following organizations have been admitted to the Institute as Company Affiliates 3M South Africa (Pty) Limited
Expectra 2004 (Pty) Ltd
Murray and Roberts Cementation
AECOM SA (Pty) Ltd
Exxaro Coal (Pty) Ltd
Nalco Africa (Pty) Ltd
AEL Mining Services Limited
Exxaro Resources Limited
Namakwa Sands(Pty) Ltd
Air Liquide (Pty) Ltd
Filtaquip (Pty) Ltd
Ncamiso Trading (Pty) Ltd
Alexander Proudfoot Africa (Pty) Ltd
FLSmidth Minerals (Pty) Ltd
New Concept Mining (Pty) Limited
AMEC Foster Wheeler
Fluor Daniel SA ( Pty) Ltd
Northam Platinum Ltd - Zondereinde
AMIRA International Africa (Pty) Ltd
Franki Africa (Pty) Ltd-JHB
Opermin Operational Excellence
ANDRITZ Delkor (Pty) Ltd
Fraser Alexander (Pty) Ltd
OPTRON (Pty) Ltd
Anglo Operations Proprietary Limited
G H H Mining Machines (Pty) Ltd
PANalytical (Pty) Ltd
Anglogold Ashanti Ltd
Geobrugg Southern Africa (Pty) Ltd
Arcus Gibb (Pty) Ltd
Glencore
Paterson & Cooke Consulting Engineers (Pty) Ltd
ASPASA
Hall Core Drilling (Pty) Ltd
Perkinelmer
Atlas Copco Holdings South Africa (Pty) Limited
Hatch (Pty) Ltd
Polysius A Division Of Thyssenkrupp Industrial Sol
Aurecon South Africa (Pty) Ltd
HPE Hydro Power Equipment (Pty) Ltd
Aveng Engineering
Immersive Technologies
Aveng Mining Shafts and Underground
IMS Engineering (Pty) Ltd
Axis House Pty Ltd
Ingwenya Mineral Processing
Bafokeng Rasimone Platinum Mine
Ivanhoe Mines SA
Barloworld Equipment -Mining
Joy Global Inc.(Africa)
BASF Holdings SA (Pty) Ltd
Kudumane Manganese Resources
BCL Limited
Leco Africa (Pty) Limited
Becker Mining (Pty) Ltd
Longyear South Africa (Pty) Ltd
BedRock Mining Support (Pty) Ltd
Lonmin Plc
BHP Billiton Energy Coal SA Ltd
Lull Storm Trading (Pty) Ltd
Blue Cube Systems (Pty) Ltd
Maccaferri SA (Pty) Ltd
Bluhm Burton Engineering Pty Ltd
Magnetech (Pty) Ltd
Bouygues Travaux Publics
Magotteaux (Pty) LTD
Sandvik Mining and Construction RSA(Pty) Ltd
CDM Group
Maptek (Pty) Ltd
SANIRE
CGG Services SA
MBE Minerals SA (Pty) Ltd
Schauenburg (Pty) Ltd
MCC Contracts (Pty) Ltd
Sebilo Resources (Pty) Ltd
MD Mineral Technologies SA (Pty) Ltd
SENET (Pty) Ltd
MDM Technical Africa (Pty) Ltd
Senmin International (Pty) Ltd
Metalock Engineering RSA (Pty) Ltd
Smec South Africa
Metorex Limited
Sound Mining Solution (Pty) Ltd
Metso Minerals (South Africa) (Pty) Ltd
Speciality Construction Products (Pty) Ltd
CSIR Natural Resources and the Environment (NRE)
Micromine Africa (Pty) Ltd
SRK Consulting SA (Pty) Ltd
Minerals Council of South Africa
Time Mining and Processing (Pty) Ltd
Data Mine SA
Minerals Operations Executive (Pty) Ltd
Timrite (Pty) Ltd
Digby Wells and Associates
MineRP Holding (Pty) Ltd
Tomra (Pty) Ltd
DRA Mineral Projects (Pty) Ltd
Mintek
Ukwazi Mining Solutions (Pty) Ltd
DTP Mining - Bouygues Construction
MIP Process Technologies (Pty) Limited
Umgeni Water
Duraset
Modular Mining Systems Africa (Pty) Ltd
Webber Wentzel
Elbroc Mining Products (Pty) Ltd
MSA Group (Pty) Ltd
Weir Minerals Africa
eThekwini Municipality
Multotec (Pty) Ltd
Worley (Pty) Ltd
Coalmin Process Technologies CC Concor Opencast Mining Concor Technicrete Council for Geoscience Library CRONIMET Mining Processing SA (Pty) Ltd
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APRIL 2019
Herrenknecht AG
VOLUME 119
Precious Metals Refiners Ramika Projects (Pty) Ltd Rand Refinery Limited Redpath Mining (South Africa) (Pty) Ltd Rocbolt Technologies Rosond (Pty) Ltd Royal Bafokeng Platinum Roytec Global (Pty) Ltd RungePincockMinarco Limited Rustenburg Platinum Mines Limited Salene Mining (Pty) Ltd Sandvik Mining and Construction Delmas (Pty) Ltd
The Journal of the Southern African Institute of Mining and Metallurgy
2019/06/24 8:24 AM
GLOBAL MINERAL PROFESSIONALS ALLIANCE Global Mineral Professionals Alliance undertakes International Action on Tailings In light of the global challenges faced around mine waste, particularly tailings, the Global Mineral Professional Alliance (GMPA) will be undertaking ‘Global Action on Tailings’. The initiative was announced at a recent meeting of the GMPA – a collaborative partnership between leading professional organizations for minerals around the world – held in February 2019, in Denver, Colorado in conjunction with the 2019 SME Annual Conference & Expo. In addition to examining tailings from the perspective of disposal and closure, the initiative will seek to advance discussions and solutions relating to tailings management, tailings reprocessing/repurposing, and ultimately moving towards tailings elimination. “Today, tailings management is the most pressing issue facing the minerals sector,” said Janice Zinck, President of the Canadian Institute of Mining, Metallurgy and Petroleum. “Collectively, the GPMA has thousands of experts in all aspects of mining and milling relating to tailings generation, reprocessing and management, making its members well poised to address this complex topic from a scientific, fact-based perspective.” “Over the next several months the GMPA organizations will collaborate on this initiative to develop working groups that can effectively address the significant challenges associated with tailings management at existing impoundments, as well as focus on the need to reduce, and ultimately eliminate, the use of conventional surface tailings storage facilities,” SME President Dr. Hugh Miller further added. “As a Society, we have tremendous technical capacity to confront these challenges through our members and fully recognize the need for global collaboration to make substantive change.” The approach will be multi-pronged, with the first step involving data collection and the dissemination of existing best practices, case studies, and research though a forthcoming GMPA Global Action on Tailings website. “IOM3’s members have significant expertise in these issues and the Institute itself holds an immense amount of research material from across the world on all aspects of tailings management,” said Prof. Serena Best, President of the Institute of Materials, Minerals and Mining. “This important initiative will help us apply this expertise and knowledge to address this significant challenge.” AusIMM President Janine Herzig endorsed this initiative saying that the most recent tailings dam disaster in Brazil was a tragedy, and the AusIMM welcomed the report from a full inquiry. “Australasia is respected as a world leader in best practice health, safety, community and environmental standards through a robust regulatory framework informed by sought-after mining health and safety experts,” Ms. Herzig said. “AusIMM continues to be committed to educating and promoting world class standards, and investing in the capability of professionals in the industry ensuring that all those who work in the resources sector continue to be safe and healthy.” Tailings teams will be established in each of the participating GMPA organizations and experts on all aspects of tailings management and processing will be engaged in addressing this issue. “The Southern African Institute of Mining and Metallurgy fully supports this important initiative. Recent events have galvanized focus on the design, placement and management of waste storage facilities, while at the same time emphasising the need for research to reduce the need for such facilities. The sharing of global knowledge in these efforts is essential, not only to pool global knowledge, but also to prevent sub-optimisation on a local basis,” stated SAIMM President, Alastair Macfarlane. To emphasise support in a Southern African context, the SAIMM is holding a Tailings Storage conference in October. A think tank event will be held in Arequipa, Peru in September, 2019 in conjunction with PERUMIN — Convencion Minera. “IIMP is delighted to host the inaugural Global Action on Tailings event in Arequipa this coming September,” said Antonio Samaniego, Past President of IIMP. The GMPA also hopes to work collaboratively with the International Council on Mining and Metals (ICMM) which recently announced its intention to establish an independent panel of experts to develop an international standard for tailings facilities for its member companies. About GMPA: Comprised of the Australasian Institute of Mining and Metallurgy (AusIMM), the Canadian Institute of Mining, Metallurgy and Petroleum (CIM), the Peruvian Institute of Mining Engineers (IIMP), the Institute of Materials, Minerals and Mining (IOM3), the Southern African Institute of Mining and Metallurgy (SAIMM), and the Society for Mining, Metallurgy & Exploration (SME), the GMPA strives to foster the exchange of knowledge and technology, promote technical and professional excellence across the minerals industries, provide programs for education and professional development, and facilitate opportunities for knowledge sharing, business networking, fellowship and public information. ###
SME is a professional society (nonprofit 501(c)(3) corporation) whose members represent all professions serving the mining, minerals and underground construction industries in more than 100 countries. SME members include engineers, geologists, metallurgists, educators, students and researchers. SME advances the worldwide minerals community through information exchange and professional development.
News Release
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SME exists to be the premier resource and advocate for the mining community. For further details, see www.smenet.org
2019/06/18 8:10 AM
Resilient futures for mineral processing We would like to invite you to attend the XXX International Mineral Processing Congress in Cape Town, South Africa from 18 to 22 October 2020. IMPC 2020 will be hosted by the Southern African Institute of Mining and Metallurgy (SAIMM).
XXX INTERNATIONAL MINERAL PROCESSING CONGRESS 18 - 22 OCTOBER
CAPE TOWN
SOUTH AFRICA
www.impc2020.com
THE SAIMM IMPC 2020 will be hosted by the Southern African Institute of Mining and Metallurgy (SAIMM). The SAIMM has been in existence for 125 years, having been established in 1894 as a ‘learned society’ to support mining and metallurgical professionals during the emergence and growth of the early South African minerals industry. Mining is of great importance to Africa in general, and particularly to Southern Africa. Africa accounts for a major portion of the world’s mineral reserves and more than half of gold, platinum group metals, cobalt and diamonds. Southern Africa produces over two-thirds of Africa’s mineral exports by value.
CAPE TOWN INTERNATIONAL CONVENTION CENTRE IMPC 2020 will be hosted at Cape Town International Convention Centre (CTICC). Since the inception of the CTICC in 2003, Cape Town has been proudly the number one destination for conferences in Africa, according to the latest International Congress and Convention Association (ICCA) statistics. Cape Town, the “Mother City”, is the oldest city in South Africa and has a cultural heritage spanning more than 300 years. Cape Town is a modern, cosmopolitan city and is often rated as one of the premier world holiday destinations. The city has a large range of hotels & guest houses and modern transport infrastructure. The city has numerous activities & attractions, including Table Mountain, Robben Island, Cape Point, the Castle, V&A Waterfront, world class beaches, wine farms, nature reserves, scenic drives, hiking, whale watching, shark cage diving and fine dining. Photo courtesy CTICC
CAPE TOWN INTERNATIONAL CONVENTION SQUARE 1 LOWER LONG STREET CAPE TOWN 8001
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2019/06/18 8:10 AM